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NUREG/IA-0211 
IRSN 2005-194 
NSI RRCKI3188 

International 
Agreement Report 


Experimental Study of Embrittlement 
of Zr-1 %Nb VVER Cladding under 
LOCA-Relevant Conditions 


Prepared by 

L. Yegorova, K. Lioutov, N. Jouravkova, A. Konobeev 

Nuclear Safety Institute of Russian Research Centre “Kurchatov Institute” 

Kurchatov Square 1, Moscow 123182, Russian Federation 

V. Smirnov, V. Chesanov, A. Goryachev 

State Research Centre “Research Institute of Atomic Reactors” 

Dimitrovgrad 433510, Russian Federation 


Office of Nuclear Regulatory Research 
U.S. Nuclear Regulatory Commission 
Washington, DC 20555-0001 

March 2005 


Prepared for 

U.S. Nuclear Regulatory Commission (United States), 

Institute for Radiological Protection and Nuclear Safety (France), 
and Joint Stock Company "TVEL" (Russian Federation) 


Published by 

U.S. Nuclear Regulatory Commission 








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IRSN 2005-194 
NSI RRC KI 3188 


International 
Agreement Report 


Experimental Study of Embrittlement 
of Zr-1 %Nb VVER Cladding under 
LOCA-Relevant Conditions 


Prepared by 

L. Yegorova, K. Lioutov, N. Jouravkova, A. Konobeev 

Nuclear Safety Institute of Russian Research Centre “Kurchatov Institute” 

Kurchatov Square 1, Moscow 123182. Russian Federation 

V. Smirnov, V. Chesanov, A. Goryachev 

State Research Centre “Research Institute of Atomic Reactors” 

Dimitrovgrad 433510. Russian Federation 


Office of Nuclear Regulatory Research 
U.S. Nuclear Regulatory Commission 
Washington, DC 20555-0001 

March 2005 


Prepared for 

U.S. Nuclear Regulatory Commission (United States), 

Institute for Radiological Protection and Nuclear Safety (France), 
and Joint Stock Company "TVEL" (Russian Federation) 


Published by 

U.S. Nuclear Regulatory Commission 








tKW 7 

,5 

% E°C\ 

l.oo5 

Opy 


i 


Zdo 6?53I 3o8 


Abstract 


During 2001-2004. research was performed to develop test data on the embrittlement of niobium-bearing 
(Zr-l%Nb) cladding of the VVER type under loss-of-coolant accident (LOCA) conditions. Procedures were 
developed and validated to determine the zero ductility threshold. Pre-test and post-test examinations in¬ 
cluded weight gain and hydrogen content measurements, preparation of metallographic samples, and exami¬ 
nation of samples using optical microscopy, scanning electron microscopy and transmission electron micros¬ 
copy. Sensitivity of the zero ductility threshold to heating and cooling rates was determined. Oxidation 
kinetics and ductility threshold were measured for the standard El 10 alloy, six variants with different impuri¬ 
ties. Zircaloy, and irradiated El 10. Oxidation temperatures were varied from 800-1200 C, and mechanical 
(ring compression) testing temperatures were varied from 20-300 C. It was concluded that (a) the current 
type of El 10 cladding has an optimal microstructure, (b) oxidation and ductility of the oxidized cladding are 
very sensitive to microchemical composition and surface finish, (c) the use of sponge zirconium for fabrica¬ 
tion of cladding tubes provides a significant reduction of the oxidation rate and an increase in the zero ductil¬ 
ity threshold, and (d) additional improvement in oxidation and ductility can be achieved by polishing the 
cladding surface. 


m 


LC Control Number 



2006 


531308 





























Foreword 


A world-wide trend to substantially increase nuclear fuel bumup to higher levels has led fuel manufacturers 
in the U.S. and France to develop niobium-bearing cladding alloys that are similar in composition to Russian 
cladding alloys. These alloys, E-110, E-635, ZIRLO, and M5, all have greatly improved corrosion resistance 
compared with Zircaloy during normal operation, especially at higher bumup levels. However, in early 
2001, it was realized that the Russian alloys and the Western niobium-bearing alloys behaved somewhat 
differently under conditions of a loss-of-coolant accident (LOCA), during which the cladding is exposed to 
steam at high temperatures. 


At that time, a research program was already underway at Argonne National Laboratory in the U.S. to inves¬ 
tigate the effects of high-bumup on cladding behavior under LOCA conditions. Further, a cooperative re¬ 
search effort on fuel behavior was also underway at the Russian Research Center (Kurchatov Institute) with 
partial sponsorship by the French Institute for Radiological Protection and Nuclear Safety and the U.S. Nu¬ 
clear Regulatory Commission; additional funding was being provided by the Russian fuel vendor, TVEL. It 
was then decided to investigate the underlying phenomena that governed cladding behavior of niobium¬ 
bearing alloys — particularly the Russian cladding — in this Russian program. By closely coordinating this 
research with the w'ork underway at Argonne National Laboratory on similar Western alloys, it was hoped 
that a fuller understanding could be obtained. Coordination between laboratories was further enhanced by 
including some Zircaloy cladding specimens in the Russian program and including some El 10 cladding 
specimens in the program at Argonne National Laboratory. 


After several years of research at both laboratories, the general cause of differences in behavior under LOCA 
conditions has been isolated to the ore reduction process and the surface finish of the cladding tubing. This 
understanding is helping to improve licensing criteria that can be applied to new and different cladding al¬ 
loys. The extensive work performed by the Russian Research Center (Kurchatov Institute) and their collabo¬ 
rating laboratory at the State Research Center (Research Institute of Atomic Reactors) is documented in the 
following report. 



Senior Technical Advisor 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 



Director 

Office of Nuclear Regulatory Research 
U.S. Nuclear Regulatory Commission 


v 




























TABLE OF CONTENTS 


Page 

1. INTRODUCTION.1.1 

2. BACKGROUND.2.1 

3. OXIDATION BEHAVIOR AND EMBRITTLEMENT THRESHOLD OF STANDARD El 10 

CLADDING: PROGRAM, TEST PROCEDURES, DISCUSSION OF TEST RESULTS.3.1 

3.1. The program concept and technical requirements to experimental works.3.1 

3.2. Methodological aspects of oxidation and mechanical tests.3.6 

3.2.1. Oxidation tests . 3.6 

3.2.2. Mechanical tests . 3.11 

3.3. Discussion of test results and working out of preliminary conclusions.3.26 

3.3.1. Reference tests . 3.26 

3.3.2. The comparative analysis of El 10 and Zry-4 oxidation and mechanical behavior . 3.29 

3.3.3. Determination of sensitivity of the El 10 cladding embrittlement to the oxidation type and the 

characterization of comparative behavior of El 10 and M5 claddings . 3.48 

3.3.4. The evaluation of the El 10 oxidation and mechanical behavior as a function of oxidation 

temperature . 3.52 

3.3.5. The sensitivity of the behavior of Russian niobium-bearing alloys to the alloying 

composition . 3.60 

3.3.6. Interrelation between the zero ductility threshold and the temperature of mechanical 

tests . 3.62 

3.3. 7. The analysis of representativitv of the zero ductility threshold determined due to ring 

compression tests . 3.68 

3.3.8. Consideration of the zero ductility threshold of the El 10 cladding as a function of the 

irradiation effect . 3.75 

3.3.9. The analysis of the El 10 oxidation kinetics . 3.80 

4. OXIDATION BEHAVIOR AND EMBRITTLEMENT THRESHOLD OF THE MODIFIED El 10 

CLADDING: PROGRAM AND DISCUSSION OF TEST RESULTS.4.1 

4.1. Major provisions of the test program with a modified E110 cladding.4.1 

4.2. The analysis of experimental results obtained at surface effect studies.4.4 

4.3. The assessment of relationship between the microchemical composition and oxidation 

BEHAVIOR OF NIOBIUM-BEARING ALLOYS .4.9 

4.3.1. The analysis of the oxidation and mechanical behavior for the El 10 Gl j rJ and E635 G( f r) 

claddings fabricated on the basis of 100% French sponge Zr . 4.10 

vii 
























4.3.2. The interpretation of test results with EllOcafr) and El 10G<3no claddings . 4.11 

4.3.3. The sensitivity of the oxidation behavior of the sponge El 10 cladding to the oxidation 

temperature . 4.13 

4.3.4. The analysis of results obtained in the test with the E635 cladding fabricated . 

using sponge Zr . 4.17 

4.3.5. The comparative consideration of a microchemical composition of different types . 

of the E110 alloy . 4.18 

4.4. The comparative analysis of the E110 material microstructure.4.24 

4.5. The oxidation kinetics of the sponge type E110 cladding.4.28 

5. SUMMARY.5.1 

5.1. Major findings of the program first part.5.1 

5.1.1. Methodological aspects of mechanical tests . 5. 1 

5.1.2. Methodological aspects of oxidation tests . 5.4 

5.1.3. The embrittlement behavior of Zr-1 %Nb (El 10) cladding . 5.5 

5.1.4. The embrittlement behavior ofZr-l%Nb (El 10) cladding as a function 

of irradiation effects . 5.8 

5.1.5. The oxidation kinetics and embrittlement behavior of the El 10 cladding according 

to results of previous investigations in different laboratories . 5. 9 

5.2. Major findings of the program second part.5.11 

5.2.1. The concept of special investigations . 5.11 

5.2.2. Surface effect studies . 5.11 

5.2.3. Bulk chemistry studies . 5.12 

5.2.4. Bulk microstructure studies . 5.16 

5.2.5. Final Remarks . 5.17 


viii 























LIST OF FIGURES 


Page 

Fig. 3.1. Schematics of the cladding sample.3.7 

Fig. 3.2. Oxidation test apparatus.3.8 

Fig. 3.3. Types of temperature histories for different combinations of heating and cooling rates.3.9 

Fig. 3.4. Development of the data base with test results.3.10 

Fig. 3.5. Test machine for ring compression tests of oxidized cladding samples.3.11 

Fig. 3.6. The as-measured load-displacement diagram of the El 10 oxidized cladding sample.3.12 

Fig. 3.7. Procedure for the determination of the sample zero-displacement point.3.13 

Fig. 3.8. Determination of effective modulus of elasticity basing on the results of scoping compression 

tests.3.14 

Fig. 3.9. The processing procedure for the load-displacement diagram of the ring compression test of 

the E110 oxidized cladding.3.15 

Fig. 3.10. Relative displacement at failure of the El 10 cladding after a double-sided oxidation and S/S 

combination of heating and cooling rates as a function of the ECR.3.15 

Fig. 3.11. The validation of the procedure for the zero-ductility threshold determination.3.16 

Fig. 3.12. Sensitivity of a relative displacement at failure to the ring sample length for the ductile 

sample.3.17 

Fig. 3.13. Sensitivity of a relative displacement at failure to the ring sample length for the brittle 

sample.3.17 

Fig. 3.14. Typical El 10 load-displacement diagrams for different oxidation.3.18 

Fig. 3.15. Demonstration of the state of ductile claddings after ring compression tests.3.19 

Fig. 3.16. The data base for the interpretation of load-displacement diagrams for oxidized samples 

with a high ductility margin.3.20 

Fig. 3.17. The load-displacement diagram of Zry-4 cladding sample with a partial ductility margin.3.21 

Fig. 3.18. The data base to characterize the mechanical behavior of the cladding sample with a partial 

residual ductility before the fracture.3.21 

Fig. 3.19. The data base to characterize the mechanical behavior of cladding sample with the partial 

residual ductility at failure.3.22 

Fig. 3.20. The data base to characterize the mechanical behavior of the brittle ring sample.3.23 

Fig. 3.21. Processing of the load-displacement diagram obtained in the ring tensile test of a simple ring 

sample manufactured from the oxidized El 10 cladding tube.3.24 


IX 























Fig. 3.22. Three-point bending test apparatus.3.25 

Fig. 3.23. Schematic for the processing of the load-displacement diagram after three-point bending 

tests.3.25 

Fig. 3.24. Results of reference tests.3.27 

Fig. 3.25. The residual ductility of the El 10 cladding vs heating and cooling rates.3.28 

Fig. 3.26. The zero ductility threshold of the El 10 cladding vs heating and cooling rates.3.28 

Fig. 3.27. Summary of the ring compression test results performed in different laboratories with the 

Zry-4 oxidized cladding.3.30 

Fig. 3.28. The comparative test data characterizing the ductility of Zry-4 oxidized cladding vs ECR.3.31 

Fig. 3.29. The comparative data obtained in different laboratories to characterize the Zry-4 oxidation 

kinetics.3.32 

Fig. 3.30. The comparison of El 10 and Zry-4 cladding behavior in accordance with RRC KI/RIAR 

test data.3.32 

Fig. 3.31. Appearance of the El 10 and Zry-4 claddings after the oxidation at 11.3-11.8% ECR and 

F/F, S/S combinations of heating and cooling rates.3.33 

Fig. 3.32. The appearance in detail of the El 10 oxidized surface vs the ECR.3.34 

Fig. 3.33. Demonstration of two layers of Zr0 2 oxide on the outer surface of the El 10 oxidized 

standard as-received tube using fractography results.3.34 

Fig. 3.34. Cross section of the oxide nodule in the El 10 cladding.3.35 

Fig. 3.35. The characterization of the oxidized El 10 claddings vs ECR.3.35 

Fig. 3.36. Visualization of Zr0 2 oxide behavior as a function of the ECR after the double-sided 

oxidation at 1100 C (F/F combination of heating and cooling rates) of El 10 standard as-received 
tubes.3.36 

Fig. 3.37. The appearance and microstructure of Zr-Nb oxidized cladding after the operation with the 

surface boiling (RBMK cladding type).3.37 

Fig. 3.38. Angular variations of the El 10 standard as-received tube microstructure after a double-sided 

oxidation at 1100 C and 10% ECR (sample #17).3.37 

Fig. 3.39. The microstructure of the Zry-4 cladding (sample #64) oxidized at 11.5% ECR at F/F 

combination of heating and cooling rates.3.38 

Fig. 3.40. The microstructure of the El 10 cladding after a single-sided oxidation at 1100 C.3.39 

Fig. 3.41. The comparative data characterizing the Zr0 2 and a-Zr(O) thickness.3.39 


x 






















Fig. 3.42. The appearance of microstructure and characterization of oxygen distribution in the El 10 

cladding after the double-sided oxidation at 1100 C.3.41 

Fig. 3.43. The oxygen distribution in the El 10 cladding (1100 C, 8.2% ECR) in accordance with 

results of SEM examinations.3.42 

Fig. 3.44. The comparison of microhardness distributions for the brittle and ductile El 10 oxidized 

cladding. 3.43 

Fig. 3.45. Demonstration of the morphology of ZrCE layers in the El 10 oxidized cladding (sample 

#41-4, 1100 C, 8.2% ECR).3.44 

Fig. 3.46. The niobium distribution in the El 10 cladding (1100 C, 8.2% ECR) in accordance with 

results of SEM examinations.3.46 

Fig. 3.47. The SEM micrograph of the El 10 oxidized cladding.3.47 

Fig. 3.48. The optical microstructure of the El 10 oxidized sample with hydrides in the prior |3-phase.3.48 

Fig. 3.49. The comparison of the El 10 appearance and microstructure after single-sided and double¬ 
sided oxidation at 1100 C.3.49 

Fig. 3.50. Comparative data characterizing the El 10 residual ductility as a function of the oxidation 

type.3.49 

Fig. 3.51. The data base characterizing the mechanical behavior of the El 10 and M5 claddings after a 

single-sided oxidation at 1100 C.3.50 

Fig. 3.52. The comparison of the El 10 and M5 cladding mechanical behavior after the single-sided 

oxidation at 1100 C in accordance with the ring compression test results at 20 C.3.51 

Fig. 3.53. The characterization of the El 10 commercial cladding after irradiation and oxidation in the 

a-phase.3.52 

Fig. 3.54. The characterization of the allotropic phase transformation in the El 10 and M5 alloys .3.53 

Fig. 3.55. Appearances of the El 10 cladding after the double-sided oxidation at 800-1100 C and F/F 

combination of heating and cooling rates.3.54 

Fig. 3.56. The microstructure of the El 10 cladding after a double-sided oxidation at 800-950 C and 

F/F combination of heating and cooling rates.3.55 

Fig. 3.57. The data base characterizing the residual ductility of the El 10 cladding as a function of the 

ECR and oxidation temperature.3.56 

Fig. 3.58. The El 10 residual ductility and hydrogen concentration as a function of the ECR after a 

double-sided oxidation at 1100 C'and F/F, F/Q combinations of heating and cooling rates.3.57 

Fig. 3.59. The hydrogen content in the El 10 cladding as a function of the ECR and temperature after a 

double-sided oxidation.3.57 


xi 




















Fig. 3.60. The comparison of the El 10 microstructure after a double-sided oxidation at different 

temperatures.3.58 

Fig. 3.61. The characterization of the El 10 cladding behavior after a double-sided oxidation at 1200 C.. 3.59 

Fig. 3.62. The summary of results characterizing the El 10K behavior under oxidation and ring 

compression test conditions.3.61 

Fig. 3.63. The characterization of the E635 cladding behavior under oxidation and mechanical test 

conditions.3.63 

Fig. 3.64. Load-displacement diagrams for two Zry-2 samples hydrided (specially) up to 180 ppm and 
700 ppm, respectively, as a function of the temperature mechanical tests of the bending type 
(reprinted from [47]).3.65 

Fig. 3.65. Dependence of the El 10 cladding ductility on the ECR (oxidation at 1100 C) and 

temperature of ring tensile tests.3.65 

Fig. 3.66. Residual ductility of two El 10 samples oxidized at 10 and 11.7% ECR (1100 C) as a 

function of temperature ring compression tests.3.66 

Fig. 3.67. The sensitivity of the El 10 residual ductility (800-1200 C, F/f and F/Q) to the hydrogen 

concentration at 20 and 135 C.3.67 

Fig. 3.68. The data characterizing the sensitivity of the El 10 residual ductility at 135 C to the ECR 

(900-1100 C, F/F and F/Q).3.68 

Fig. 3.69. The comparison of zero ductility thresholds determined from the results of ring tensile and 

ring compression tests (El 10, 1100 C).3.69 

Fig. 3.70. The zero ductility threshold of the El 10 cladding determined due to three-point bending 

tests (1100 C, F/F).3.69 

Fig. 3.71. SEM micrographs for fracture surfaces of the El 10 brittle cladding.3.70 

Fig. 3.72. High magnification SEM micrographs of fracture surface regions of the El 10 brittle 

cladding.3.71 

Fig. 3.73. The SEM micrograph for the fracture surface of the El 10 ductile sample.3.71 

Fig. 3.74. The maximum load on the El 10 oxidized sample as a function of residual ductility.3.72 

Fig. 3.75. The comparative data characterizing the El 10 residual ductility as a function of the ECR 

obtained on the processing of test data of different laboratories.3.73 

Fig. 3.76. Hydrogen distribution along the width of a special Zry-4 sample.3.74 

Fig. 3.77. Demonstration of the end effects on the El 10 oxidized cladding samples.3.75 

Fig. 3.78. The appearance and microstructure of the El 10 irradiated cladding before the tests and after 

the oxidation tests at 1100 C and F/F combination of heating and cooling rates.3.76 

xii 




















Fig. 3.79. The appearance and microstructure of the El 10 oxidized cladding after the slow transient 

oxidation at 1100 C and standard oxidation at 1200 C.3.77 

Fig. 3.80. The residual ductility of the El 10 irradiated cladding as a function of the ECR.3.78 

Fig. 3.81. Comparative data characterizing the microhardness of the El 10 oxidized irradiated 

claddings.3.78 

Fig. 3.82. The zero ductility threshold of the El 10 irradiated cladding after different oxidation modes ... 3.79 

Fig. 3.83. The El 10 hydrogen concentration as a function of irradiation and ECR. residual ductility of 

the E110 irradiated cladding as a function of hydrogen concentration.3.80 

Fig. 3.84. Determination of the rate constant at the oxidation of the El 10 unirradiated cladding in the 

temperature range of 1073-1473 K.3.81 

Fig. 3.85. Comparison of the oxidation kinetics for the El 10 unirradiated and irradiated claddings.3.82 

Fig. 3.86. The comparison of data characterizing the transient test modes with the El 10 oxidation 

kinetics.3.83 

Fig. 3.87. The comparative data on the El 10 and E635 oxidation kinetics.3.83 

Fig. 3.88. The comparison of the El 10 and Zry-4 oxidation kinetics.3.84 

Fig. 3.89. The El 10 oxidation kinetics at 1100 C in accordance with the data obtained in different 

laboratories.3.84 

Fig. 4.1. The relationship between surface scratches and localization of the breakaway oxidation areas .. 4.2 

Fig. 4.2. The appearance of the El 10 claddings fabricated with the use of two different types of 

surface finishing before and after oxidation tests at 1100 C.4.5 

Fig. 4.3. The microstructure of the El 10A and El 10 m claddings after the oxidation at 1100 C.4.5 

Fig. 4.4. The comparison of residual ductility of the standard El 10 as-received tube, El 10A and 

Ell 0 m claddings after the oxidation at 1100 C.4.6 

Fig. 4.5. Demonstration of sensitivity of the oxidation and mechanical behavior for the El 10 cladding 

to the cladding surface polishing.4.7 

Fig. 4.6. The oxide microstructure on the polished and unpolished parts of the El 10 cladding after the 

oxidation at 1000 C.4.8 

Fig. 4.7. Appearances of E11 O^fr) samples after the double-sided oxidation at 1100 C.4.10 

Fig. 4.8. Results of ring compression tests with El lO^e-i oxidized samples.4.11 

Fig. 4.9. Comparative data base characterizing the El 10 G (3fr» and El 10 G( 3 m) oxidation/mechanical 

behavior.4.12 


xm 




















Fig. 4.10. The comparison of microstructures for iodide/elcctrolytic and sponge El 10 claddings after 

the oxidation at 1100 C.4.12 

Fig. 4.11. The appearance of the El 10| OwH r cladding after the oxidation at 1100 C and results of 

mechanical tests.4.13 

Fig. 4.12. The appearance and mechanical properties of different El 10 claddings after the oxidation at 

1000 C.4.15 

Fig. 4.13. The comparative test data characterizing the El 10 and El 10 G(3ru) behavior after the oxidation 

at 900 C.4.16 

Fig. 4.14. The characterization of appearance and RT mechanical properties of the El 10 G (3ru) cladding 

after the oxidation at 1200 C.4.17 

Fig. 4.15. The appearance of the E635 G(fr ) cladding after the oxidation test at 1100 C and comparative 

E635 (standard), El 10 (standard), E635 G( f r) results of ring compression tests.4.18 

Fig. 4.16. Comparison of some data on the impurity content in the standard El 10 and El 10 G(fr) , 

El 10 G( 3 f r) , El 10 G(3ru) at the beginning of the cladding fabrication.4.21 

Fig. 4.17. Low magnification of TEM micrographs for El 10, El 10 G(fr ), El 10| OW Ht claddings and the 

M5 cladding [reprinted from 27].4.25 

Fig. 4.18. The characterization of the SPP distribution in the a-Zr grains of the El 10 and El 10 G( f r) 

claddings.4.25 

Fig. 4.19. The SPP distribution in the El 10 and El 10 G(fr ) claddings.4.26 

Fig. 4.20. High magnification of the SPP TEM micrograph in the El 10 G(fr) cladding.4.27 

Fig. 4.21. The oxidation kinetics of different types of Zr-l%Nb claddings at 1100 and 1200 C.4.29 

Fig. 4.22. The oxidation kinetics of different types of Zr-l%Nb claddings at 900 and 1000 C.4.30 

Fig. 4.23. The characterization of the oxidation rate for the standard El 10 (iodide/electrolytic) and 

E110 claddings manufactured with the use of sponge Zr in the temperature range 900-1200 C.4.31 

Fig. 4.24. The difference in the thicknesses of Zr0 2 and a-Zr(O) layers in the El 10 cladding of sponge 

and iodide/electrolytic types at the oxidation at 1000 C.4.32 

Fig. 4.25. The difference in the formation of oxide and a-Zr(O) layers in the El 10 cladding of sponge 

type at 1000 and 1100 C.4.32 

Fig. 5.1. Outline of the research program.5.1 

Fig. 5.2. The schematics of ring compression tests.5.2 

Fig. 5.3. The schematic of previous approach to the determination of zero ductility threshold.5.3 

Fig. 5.4. The interpretation of ring compression test results on the basis of the load-displacement 

diagram.5.3 


xiv 






















Fig. 5.5. The variability of oxidation tests with different heating and cooling rates.5.4 

Fig. 5.6. The appearance of the El 10 and Zry-4 claddings (1100 C) as a function of the ECR.5.5 

Fig. 5.7. The El 10 residual ductility and hydrogen concentration as a function of the ECR after the 

double-sided oxidation at 1100 C and F/F, F/'Q combinations of heating and cooling rates.5.6 

Fig. 5.8. Demonstration of the El 10 breakaway oxidation effects at 950 C.5.8 

Fig. 5.9. The appearance and microstructure of the El 10 irradiated claddings as a function of the ECR .. 5.9 

Fig. 5.10 The comparison of the El 10 oxidation kinetics (1073-1473 K) in accordance with the data 

of different investigations.5.10 

Fig. 5.11. The comparison of the El 10 and Zry-4 conservative and as-measured kinetics at 1100 C.5.11 

Fig. 5.12. The appearance of the El 10 polished and unpolished parts after the oxidation at 1000 C.5.12 

Fig. 5.13. The comparison of the iodide/electrolytic and sponge El 10 cladding oxidation behavior at 

1100C.5.13 

Fig. 5.14. The comparison of the Zry-4, sponge El 10, iodide/electrolytic El 10, M5 oxidation kinetics 

at 1000 C.5.13 

Fig. 5.15. The comparison of the Zry-4, sponge El 10, iodide/electrolytic El 10 oxidation rates in the 

temperature range of 900-1200 C.5.14 


xv 





























LIST OF TABLES 


Page 

Table 2.1. The list of major investigations performed during 1974-1990 to study the fragmentation 

threshold of unirradiated Zircaloy claddings.2.3 

Table 2.2. Results of Russian investigations performed during 1980-2001 to study the oxidation and 

mechanical behavior of unirradiated Zr-l%Nb (El 10) claddings.2.6 

Table 2.3. Major results of mechanical tests with the El 10 unirradiated oxidized claddings performed 

in Germany, Hungary and Czech Republic during 1990-2000. 2.9 

Table 3.1. The list of tasks and technical requirements.3.3 

Table 4.1. The El 10 surface and bulk effects studies: major provisions of experimental program.4.4 

Table 4.2. The specification for the used cladding material.4.9 

Table 4.3. The subprogram major tasks for the test with the sponge cladding material and El 10 

cladding with low Hf content.4.9 

Table 4.4. Chemical composition of the El 10 alloy (standard).4.20 

Table 4.5. Composition of zirconium alloys used in reactor fuel design.4.21 

Table 4.6. The organized results of oxidation and mechanical tests with seven types of the standard 

E110 and modified E110.4.22 

Table 4.7. Chemical composition of the Zirlo cladding tube.4.23 

Table 4.8. Zr, Nb content in the matrix, grain boundary and P-Nb precipitates.4.26 

Table 4.9. The comparative data characterizing the microstructure of El 10, El 1 0 G( fr) claddings and the 

M5 cladding.4.27 


xvn 


















































LIST OF APPENDICES 


Page 

Appendix A. Description of Test Apparatus and Test Procedures.A-l 

Appendix B. Tables with Results of Oxidation and Mechanical Tests: El 10, El 10A, El 10K, El lOpol, 
E635, El 10G(fr), El 10G(3ru), El 10G(3fr), El lOiowHf, E635G(fr). Zry-4 as-Received Tubes and El 10 
Commercial Irradiated Claddings.B-l 

Appendix C. Temperature Histories, Appearances and Microstructures of El 10 Standard As-received 
Tubes after a Double-sided Oxidation at 1100 C and S/S, S/F, F/S Combinations of Heating and 
Cooling Rates.C-l 

Appendix D. Temperature Histories, Appearances and Microstructures of El 10 Standard As-received 
Tubes after a Double-sided Oxidation at 800, 900. 950, 1000, 1100, 1200 C and F/F (F/'Q) 
Combinations of Heating and Cooling Rates.D-l 

Appendix E. Appearance and Microstructure of El 10 Standard As-received Tubes after a Single-sided 

Oxidation at 1100 C and F/F Combination of Heating and Cooling Rates.E-l 

Appendix F. Appearances and Microstructures of E635 Standard As-received Tubes after a Double¬ 
sided Oxidation at 1000, 1100 C and F/F Combination of Heating and Cooling Rates.F-l 

Appendix G. Appearances and Microstructures of Zry-4 As-received Claddings after a Double-sided 

Oxidation at 1100 C and S/S, F/F Combinations of Heating and Cooling Rates.G-l 

Appendix H. Appearance and Microstructure of El 10, E635 As-received Tubes Manufactured on the 
Basis of the Sponge Zr, El 10iow Hf As-received Tubes after a Double-sided Oxidation at 900. 

1000, 1100, 1200 C and F/F Combination of Heating and Cooling Rates.H-l 

Appendix I. Appearance and Microstructure of El 10 Commercial Irradiated Cladding after a Double¬ 
sided Oxidation at 1000, 1100. 1200 C and S/S, S/F, F/F Combinations of Heating and Cooling 
Rates.1-1 


xix 































































Acknowledgements 


The authors ot the report are pleased to acknowledge that this research program was initiated and imple¬ 
mented with the intellectual support of the following key persons from the sponsoring organizations: P. 
Lavrenyuk (Joint Stock Company “TVEL", Russia), V. Asmolov (Russian Research Centre “Kurchatov In¬ 
stitute", Russia), R.O. Meyer (U.S. Nuclear Regulatory Commission, USA), J. Papin and G. Hache (Institute 
for Radiological Protection and Nuclear Safety, France). 

We express our profound appreciation to these persons as due to their efforts the motivated balance was 
achieved between the investigations performed in width and in depth. 

The major goal of this research was to understand the effect of niobium on the embrittlement of zirconium- 
based alloys and in particular the unique embrittlement behavior of the El 10 alloy under LOCA conditions. 
Achievement of this goal would have been impossible without the appropriate test data base developed in the 
investigations performed by the following specialists: 

• Kosvintsev Yu. Yu., Leshchenko A. Yu. (State Research Centre “Research Institute of Atomic Reactors”, 
Russia): development of the oxidation facility and of special computer system for the test data measure¬ 
ment and processing; 

• Prokhorov V.I., Makarov O.Yu. (State Research Centre “Research Institute of Atomic Reactors”, Rus¬ 
sia): performance of mechanical tests of oxidized claddings; 

• Shishalova G.V. (State Research Centre “Research Institute of Atomic Reactors”, Russia): hydrogen 
content measurements; 

• Stupina L.N., Svyatkin A.M., Zvir E.A., Ivanova I.A. (State Research Centre “Research Institute of 
Atomic Reactors”, Russia): metallographic examinations and computer processing of the cladding im¬ 
ages; 

• Novoselov A.E., Shishin V.Yu., Ostrovskiy Z.E., Yakovlev V.V., Kuzmin S.V., Obukhov A.V. (State 
Research Centre “Research Institute of Atomic Reactors”, Russia): SEM and TEM examinations of the 
cladding material; 

• Panchenko V.L., Averin S.A. (Institute of Reactor Materials, Russia): TEM examinations of the cladding 
microstructure; 

• Pylev S.S., Shestopalov A.A., Abyshov G.N. (Russia Research Centre “Kurchatov Institute”, Russia): 
analytical and calculation investigations, development of the computer data base; 

• Kaplar Ye.P. (Russian Resarch Centre “Kurchatov Institute”, Russia): development of the program prin¬ 
ciples and the preliminary analysis of test results during the research first stage. 

The authors express their sincere gratitude to all these specialists for their outstanding contribution into this 
work. Besides, the authors thank individually the above listed specialists for the permission to use the results 
of their special investigations in this report. 

Many scientific and practical aspects of this work were clarified by discussions and the exchange of views 
with several experts involved in this problem. In this connection, we would like to express our profound 
appreciation to the following experts: 

• Pimenov Yu.V. (Joint Stock Company “TVEL”, Russia); 

• Bibilashvili Yu.K., Nikulina A.V., Novikov V.V. (All-Russian Institute of Inorganic Materials - 
VNIINM - Russia); 

• Billone M., Chung H. (Argonne National Laboratory, USA). 


xxi 
























































1. Introduction 


To study the oxidation behavior and embrittlement threshold of Zr-l%Nb cladding under loss-of-coolant 
accident (LOCA) conditions, a research program was developed and implemented by the Russian Research 
Center “Kurchatov Institute” (RRC KI) in cooperation with the Research Institute of Atomic Reactors 
(RIAR). The program was performed during the period of 2001-2004 and sponsored by (a) Joint Stock 
Company “TVEL" (JSC “TVEL”, Russia), (b) U.S. Nuclear Regulatory Commission (U.S. NRC, USA), and 
(c) Institute for the Radiological Protection and Nuclear Safety (IRSN, France). 

The incentive to begin this work was directly related the increase of fuel bumup in light-water reactors 
(LWRs) to 60-70 MW d/kg U and higher. Substantial corrosion is experienced with Zircaloy claddings at 
fuel bumup higher than 50 MW d/kg U, whereas much less corrosion occurs with Zr-l%Nb cladding during 
the commercial operation in the Russian type of pressurized-water reactors (VVERs) and with niobium¬ 
bearing claddings manufactured from the M5 and Zirlo alloys after operations in the pressurized-water 
reactors (PWRs). Experimental studies performed with the VVER type of Zr-l%Nb claddings refabricated 
from commercial fuel rods with bumup up to 60 MW d/kg U have shown that this cladding has a high safety 
margin under reactivity-initiated accident (RIA) conditions. But the preliminary consideration of safety 
aspects associated with niobium-bearing claddings under LOCA conditions raised the following issues: 

• several investigations performed with Zr-l%Nb cladding of the VVER type in different countries in the 
1990s have shown that the niobium-bearing cladding has somewhat different oxidation and 
embrittlement behavior in comparison with the zircaloy cladding; 

• the same general requirements concerning the prevention of the embrittled cladding fragmentation are 
applied in the LOCA safety analysis of the VVER and PWR reactors, but different approaches are used 
for this goal. 

Taking into account these and other issues, it was decided to perform a special research program including 
the following main parts of investigations: 

• the reassessment of published data concerning the PWR and VVER cladding embrittlement under LOCA 
conditions; 

• the development and validation of test apparatus and test procedures; 

• the performance of sensitivity studies and the determination of key factors which must be studied during 
this program; 

• the performance of oxidation, mechanical tests and different pre-test and post-test examinations; 

• the analysis and interpretation of obtained results. 

The major focus of investigations performed in the frame of this work was concentrated on the 
characterization of Zr-l%Nb (El 10) oxidation and mechanical behavior as a function of such parameters as: 

• oxidation conditions (single-sided or double-sided, heating and cooling rates, oxidation temperatures 
from 800-1200 C, and weight gain); 

• mechanical test conditions (ring tensile tests, ring compression tests, three-point bending tests) and test 
temperature (20-300 C); 

• cladding irradiation (as-received and refabricated claddings from the commercial fuel rods with the 
bumup about 50 MW d/kg U); 

• cladding surface conditions (as-received tubes, as-received claddings, polished as-received tubes, ground 
as-received tubes); 

• impurity compositions in the cladding. 

In addition to oxidation and mechanical tests with Zr-l%Nb (El 10) cladding, several reference tests were 
performed with the Zircaloy-4 (zirconium-tin) and E635 (zirconium-niobium-tin) claddings. The research 
program results are presented in this report. 


1.1 















































2. Background 


The prevention of cladding fragmentation in LWR-type reactors under the LOCA conditions is one of the 
basic principles of the current safety philosophy. The reason for this is related to the following physical phe¬ 
nomena: 

• significant increase of the cladding temperature during the LOCA accident caused by the coolant 
blowdown and degradation of heat transfer; 

• high temperature cladding steam oxidation accompanied by cladding embrittlement; 

• possible fracture of the embrittled cladding caused by post-LOCA forces during quenching and post¬ 
quenching actions. 

In accordance with the current world practice, the prevention of the cladding fragmentation under LOCA 
conditions is assured by special safety criteria. Thus, the main Russian regulatory document contains two 
special requirements concerning this problem: for the Zr-l%Nb (El 10) cladding [1]: 

• the peak cladding temperature (PCT) must not exceed 1200 C; 

• the local oxidation depth (Equivalent Cladding Reacted layer, ECR) must not exceed 18% of the initial 
wall thickness. 

Similar criteria are contained in the regulatory documents of other countries for the zircaloy cladding 
(1200 C, 15-17% ECR). It should be noted that the concept for the use of fragmentation criteria of these 
types was developed by the NRC (USA) with respect to the zircaloy claddings [2]. The motivation for the 
choice of this approach for the safety fragmentation criteria was reconstructed in the recent paper prepared 
by H.Chung and G.Hache [3], 

The first research to determine the zero ductility threshold of oxidized Zry-4 claddings after the quench 
cooling was performed by D.O. Hobson at the beginning of 1970s [4, 5], In accordance with results of a slow 
compression of Zry-4 oxidized samples, he revealed the relationship between the critical relative thickness of 
the prior P-phase and zero ductility threshold at 135 C (the saturation temperature during the reflood stage). 
This relationship was used to develop the following embrittlement criterion [6]: 

A. < 0.44, 

K 

where £ T - the thickness of the oxygen-rich cladding layers (ZrCL and a-ZrO); 

W 0 - the initial thickness of the unoxidized cladding. 

Further, it was revealed that the cladding ductility margin at low temperatures (150 C or less) was a function 
of not only the prior p-phase thickness but also of the oxygen concentration in this layer. The appropriate 
analysis of Hobson’s test has shown that: 

• the maximum oxygen concentration in the prior p-phase is a function of the oxygen solubility in the 
p-phase under high temperature oxidation conditions; 

• the zero ductility threshold (at 20 - 150 C) is associated with 0.7% (by weight) oxygen concentration in 
the prior p-phase. This critical oxygen concentration is achieved very fast if the oxidation temperature is 
higher than 1204 C (2200 F). 

Taking into account these test results, the embrittlement criterion was added with the Peak Cladding Tem¬ 
perature (PCT) criterion. The PCT criterion was estimated as 1204 C [6]. 

The final evolution of these criteria involved the following: 

• the extension of the test data base needed to validate criteria [7, 8, 9]; 

• the introduction of the reasonable conservative principle into the safety analysis procedure. 

As for the conservatism, it was decided to organize the results of different tests into the unified system using 
the Baker-Just equation allowing to calculate the Zry-4 high temperature kinetics with the conservative mar¬ 
gin [10]. This equation was used to determine the oxygen weight uptake with each tested sample. Besides, to 

2.1 



improve the physical interpretation of calculated results, the concept of the Equivalent Cladding Reacted 
(ECR) layer was introduced: 

ECR = , 

S. 

where S t - oxidation equivalent layer determined using the following condition: the all uptaken oxy¬ 
gen are used for the formation of the stoichiometric zirconium dioxide (Zr 02 ); 

S 0 - the initial cladding thickness. 

The practical implementation of this approach allowed the development of a test data base with the following 
list of test parameters: 

• ECR calculated with the Baker-Just; 

• oxidation time and temperature; 

• tested cladding sample characterization: intact, failed. 

The analysis of this data base presented in Reference 3 shows that: 

• at an oxidation temperature less than 1204 C, the brittle fracture of oxidized claddings was not observed 
under ring compression test conditions (at 135 C and higher) if the ECR calculated with the Baker-Just 
correlation was less than 17%; 

• at an oxidation temperature less than 1600 C, the brittle fracture of oxidation claddings was not observed 
under thermal-shock during direct quenching conditions if the ECR calculated with the Baker-Just 
correlation was less than 19%. 

Thus, the results of two different types of tests (the comparison of mechanical tests and thermal-shock tests) 
demonstrated the similarity in the evaluation of the critical oxidized thickness (17-19% ECR) although there 
was a significant discrepancy in the estimation of the permissible peak cladding temperature. These results 
led the experts responsible for the development of the proposal on safety criteria to the formulation of their 
position concerning the choice of the permissible peak temperature [11]. This position may be characterized 
by the following general provisions: 

• the practical application of thermal-shock test results requires such a detailed knowledge of physical 
processes during the LOCA which cannot be provided; 

• to prevent the fragmentation, the oxidized cladding must retain some margin of ductility; 

• the choice of 1204 C (2200 F) PCT limit based on results of compression mechanical tests provides the 
conservative margin in comparison with the thermal-shock test results. 

In 1973, this position was used to formulate the NRC criteria [2]: 1204 C, 17% ECR (calculated using the 
Baker-Just equation). During the period of 1974-1990, experimental investigations with zircaloy claddings 
were continued to understand the sensitivity of fragmentation threshold to such factors as: 

• the cladding ballooning and burst; 

• the mechanical interaction of the oxidized cladding in the fuel bundle caused by ballooning and bending; 

• the axial mechanical constraint of ballooned cladding by the grid spacers. 

A brief description of this cycle of Zry-4 investigations is presented in Table 2.1. The main outcomes of 
these investigations may be characterized in the following way: 

1. The thermal-shock tests performed with the original geometry of the oxidized cladding (without axial 
constraint or ballooning and burst) or with the original geometry of fuel rod simulators have shown that 
the current safety criteria (1204 C, 17% ECR) have a margin of about 100% in ECR. The margin in PCT 
does not exceed 150 C. 

2. Thermal-shock studies performed to check the constraining effect of the grids when using deformed 
cladding (ballooning and burst) have shown that the fragmentation threshold decreased significantly with 
axial constraint, but the test fragmentation threshold did not exceed the safety criteria. 


2.2 


3. Special impact tests performed with the simulation of potential impact fracture energy (estimated as 
0.3 J) have shown that: 

• the oxidized cladding with the original geometry had a high margin before fracture; 

• the fragmentation threshold of the deformed cladding (after ballooning and burst) was in agreement with 
the safety criteria. Some claddings were fragmented at values lower than 17% (safety criterion), and the 
increase in hydrogen content in local parts of the oxidized claddings (local cladding hydriding) was the 
cause of this effect, but the peak cladding temperature was higher than 1204 C (safety criterion). 

4. Additional studies of the hydriding effect performed with compression tests demonstrated that the zero 
ductility threshold of hydrided cladding corresponded to 700 ppm of hydrogen in the prior P-phase. 

5. The axial tensile and ring compression tests confirmed that the zero ductility threshold was reached when 
the average oxygen concentration in the prior P-phase increased from 0.6% by weight up to 0.8% by 
weight. 

Taking into account results of all these tests and understanding of the fact that the ductility margin of the 
oxidized cladding is a function of oxygen and hydrogen concentration in the prior P phase, several investiga¬ 
tors proposed to change the current safety criteria (1204 C, 17% ECR) into criteria based on the oxygen and 
hydrogen concentrations in the oxidized claddings. But these suggestions were not apparently implemented 
due to the fact that: 

• the introduction of these criteria must be accompanied by the use of computer codes for the safety 
analysis which are able to calculate the appropriate parameters with the required accuracy; 

• numerous additional experimental programs would be needed to develop and validate the integral 
criterion based on the oxygen and hydrogen content in the prior P-phase of the oxidized cladding. 

In accordance with these considerations, improvement of the safety criteria was postponed for the time being. 

Table 2.1. The list of major investigations performed during 1974-1990 to study the fragmentation 


threshold of unirradiated Zircaloy claddings 


Test type 

Test characterization 

Test results 

1. Thermal shock tests per¬ 
formed by H.Chung and 
T.Kassner, USA, 1980 
[12] 

The double-sided oxidation of Zry-4 clad¬ 
dings at 1000-1500 C with slow cooling 
though phase transition followed by 
quench type cooling 

The cladding fragmentation threshold 
was 28% ECR (measured) at 1500 C 
and 33% ECR at 1200 C 

2. Thermal shock tests per¬ 
formed by H.Uetsuka et. 
al., Japan, 1983 [13] 

The double-sided oxidation of Zry-4 clad- 
dings at 950-1350 C with the quench 
cooling 

The cladding fragmentation was not 
noted in this temperature range at 17% 
ECR (as-calculated using Baker-Just 
correlation) 

The cladding fragmentation threshold 
was higher than 35% ECR (as- 
calculated using Baker-Just correlation) 
at 1200 C 

3. Thermal shock tests with 
the cladding axial me¬ 
chanical constrain per¬ 
formed by H.Uetsuka et. 
ah, Japan, 1983 [13] 

The double-sided oxidation of Zry-4 clad¬ 
ding with the strong fixation of one clad¬ 
ding end at 900-1300 C. The fixation of 
the second cladding end at the beginning 
of the quench cooling 

The cladding fragmentation was not 
observed in this temperature range at 
17% ECR (as-calculated using Baker- 
Just correlation) 

The cladding fragmentation threshold 
was 20% ECR (as-calculated using 
Baker-Just correlation) at 1200 C 


2.3 










Test type 

Test characterization 

Test results 

4. The in-pile tests 

(PHEBUS research reac¬ 
tor) of the tests fuel bun¬ 
dle performed by 
M.Reocreux and E. Scott 
de Martinville, France, 

1990 [14] 

The LOCA-type test (#219) without the 
quench mode with Zry-4 claddings: the 
cladding ballooning and burst at high 
temperature, the cladding oxidation 

The cladding fragmentation of one fuel 
rod occurred at the following parame¬ 
ters: 

• 16% ECR 

• 1330C 

The reason of the fragmentation was 
assessed as the mechanical constrain of 
the temperature-induced cladding re¬ 
placement 

5. Impact tests of oxidized 
claddings of original ge¬ 
ometry performed by 

H.Chung and T.Kassner, 
USA, 1980 [121 

The double-sided oxidation of Zry-4 clad¬ 
dings at 1100-1400 C and the cooling rate 

5 C/s. The impact tests of oxidized clad¬ 
dings 

The failure impact energy was higher 
than 0.8 J* [15], if the oxidation tem¬ 
perature did not exceed 1315 C and the 
ECR did not exceed 17% (as-measured 
using the metallographic method) 

6. Impact tests of deformed 
(after the ballooning and 
burst) and oxidized clad¬ 
dings performed by 
H.Chung and T.Kassner, 
USA, 1980 [12] 

Fuel rods with the Zry-4 cladding and fuel 
pellet simulators were pressurized at the 
high temperature up to the ballooning and 
burst. After that, fuel rods were oxidized 
(under isothermal conditions) and 
quenched. The impact tests of these fuel 
rods were performed at the fixed impact 
energy 0.3 J 

A good correlation between the cladding 
fragmentation threshold and 17% ECR 
(as-measured with the use of metal¬ 
lographic method), 1204 C was ob¬ 
served for the most tested fuel rods. 

A special analysis has shown that: 

• some parts of the inner surface of 
deformed claddings are character¬ 
ized by the formation of a thick 
spalled oxide 

• it is revealed that stagnant steam/ 
water conditions are responsible for 
the initiation of the breakaway oxi¬ 
dation on these parts of the cladding 

• the high hydrogen uptake up to 
2200 ppm was noted in these spe¬ 
cific zones 

• it is observed that the cladding 
fragmentation threshold sharply de¬ 
creases at the cladding hydrogen 
content 700 ppm and higher 

7. The compression tests of 
oxidized and hydrating 
claddings performed by 
H.Chung and T.Kassner, 
USA, 1980 [121 

The mechanical compression tests were 
performed with fuel rods (Zry-4 cladding, 
fuel pellet simulators) after the pressuriza¬ 
tion of the fuel cladding up to the burst 
and high temperature oxidation 

These scoping tests have shown that the 
residual ductility of the oxidized clad¬ 
ding is a strong function of oxygen and 
hydrogen content in the prior (3-phase 

8. The ring compression tests 
of deformed and oxidized 
claddings performed by 
H.Uetsuka et. al., Japan 
1981-1982 [16, 17] 

Zry-4 claddings were heated, pressurized 
up to the burst, oxidized and quenched. 
The compression tests were performed at 
100C with ring samples which were cut 
off from oxidized claddings 

A strong correlation between the hydro¬ 
gen content in the prior P-phase and 
cladding residual ductility was devel¬ 
oped 

It was shown that the cladding was fully 
embrittled at 700 ppm of hydrogen con¬ 
tent 


* The expert estimations of possible loads in the commercial fuel bundle during late stages of LOCA have shown that 
the impact energy may achieve 0.3 J 


2.4 














Test type 

Test characterization 

Test results 

9. The tensile mechanical 
tests of the oxidized per¬ 
formed by A.Sawatzky. 

UK, 1978 [18] 

The Zircaloy claddings were oxidized in 
the water steam at 1000-1600 c. After 
that, the tensile tests were performed at 
room temperature, the test data base was 
added with the microhardness measure¬ 
ments across the cladding thickness 

It was revealed that the oxygen in the 
prior f-phase was distributed nonuni- 
formely 

It was shown that the oxidized cladding 
had some ductility margin if the average 
oxygen content in the prior P-phase did 
not exceed 0.6% by weight 

It was determined that the zero ductility 
threshold of the oxidized cladding corre¬ 
sponded to 0.8% averaged oxygen con¬ 
tent (the appropriate ECR was 16%) 

10. The in-pile tests of fuel 
rods performed in the PBF 
research reactor. USA, 

1982 [19] 

The single pressurized fuel rods with the 
Zry-4 cladding were oxidized at the tem¬ 
perature transient mode 

Several fuel rods which failed after the 
tests at the post-test manipulations had 
the following parameters: 

• the fuel rod with the burst at 1100 C 
oxidized up to 12% ECR (as- 
measured) at the equivalent tem¬ 
perature 1262 C 

• three fuel rods oxidized at equiva¬ 
lent temperature 1300 C up to 5- 
11% ECR (as-measured) 

• fuel rods oxidized at equivalent 
temperature up to 7-8% ECR (as 
measured) were not fragmented 


New attempts to resume this activity were made in the middle of 1990s in the context of the increase in fuel 
bumup up to 60 MW d/kg U and higher in the LWRs. An important aspect of this new stage of investigations 
was connected with the fact that zircaloy cladding has a tendency towards the breakaway oxidation and clad¬ 
ding hydriding at the high bumup (55 MW d/kg U and higher) under normal operation conditions. These 
effects result in a decrease of cladding ductility and lead to questions concerning the mechanical behavior of 
these claddings under accident conditions. The importance of this problem was shown practically in the ex¬ 
periments with the Zry-4 irradiated cladding under reactivity-initiated accident (RIA) conditions [20, 21, 22]. 

Taking into account the revealed problems, extended investigations were initiated to study the irradiation 
effects in zircaloy claddings under LOCA conditions [23, 24, 25, 26, 27]. The important direction of these 
investigations was connected with advanced cladding materials including the niobium-bearing alloys. In this 
context, it should be noted that a niobium-bearing alloy El 10 (Zr-l%Nb) had been used as the VVER clad¬ 
ding material for several decades. Moreover, special investigations performed in 1990s showed that this alloy 
demonstrated a very high corrosion resistance under normal operation conditions up to 60 MW d/kg U [28]. 

Further, special investigations devoted to measurements of mechanical properties of El 10 irradiated clad¬ 
dings under accident conditions [29, 30, 31, 32] and experimental studies of VVER high bumup fuel rods 
(50-60 MW d/kg U) under RIA conditions [30, 33, 34, 35] have demonstrated that fuel rods with the Zr- 
l%Nb (El 10) cladding have good prospects in respect to the increase of the fuel bumup in VVERs. But it is 
obvious that the final analysis of this situation cannot be made without the consideration of experimental 
results characterizing the ductility margin of El 10 claddings after the high temperature oxidation and 
quenching under LOCA-relevant conditions. 

The official history of such investigations with the El 10 alloy was initiated in Russia goes back to the begin¬ 
ning of the 1980s. Previous investigations used for the development of the second design limit of fuel rod 
damage (1200 C, 18% ECR) in the first national regulatory document on LWR safety issues [36] were not 
described in the open publications. The outline of main Russian research programs performed during the 
1990s to develop an experimental data base characterizing the oxidation and mechanical behavior of the 
E110 cladding is presented in Table 2.2. 


2.5 









Table 2.2. Results of Russian investigations performed during 1980-2001 to study the oxidation and me¬ 


chanical behavior of unirradiated Zr-l%Nb (El 10) claddings 


Test type 

Test characterization 

Test results 

1. High temperature 
oxidation tests 
performed in 

VNIINM* and VTE** 
[37, 38, 39, 40] 

The double-sided oxidation of El 10 
claddings at 700-1500 C under 
isothermal and transient conditions 

The test data base needed to develop the El 10 
oxidation kinetics correlations was obtained 

The effect of Zr0 2 spallation was revealed in the 
temperature range 900-1100 C on achieving the 
critical oxide thickness (~25 pm) 

A significant change of the cladding geometrical 
sizes was revealed for some transient modes 

2. Ring tensile tests of 
oxidized claddings 
performed in 

VNIINM, 1990 [41] 

The double-sided oxidation of El 10 
claddings at 1000-1200 C with the 
direct current heating and fast air 
cooling. Ring tensile tests of oxidized 
samples at 20-1000 C 

The zero ductility threshold was associated with 
the weight gain 450 mg/cm' (~6% as-measured 
ECR) at 20 C 

The reasonable margin of residual ductility in 
fully brittle samples was revealed at temperatures 
higher than 500 C 

3. Impact tests of 
oxidized claddings 
performed in 

VNIINM, 1990 [41] 

The double-sided oxidation of El 10 
claddings at 1000-1200 C with the 
direct current heating and cooling rate 
10-20 C/s. The impact tests of 
oxidized claddings at 20 C 

The brittle fracture occurred at the weight gain 
higher than 600 mg/dm 2 (7% ECR) 

The relationship between the failure specific 
impact energy of the unoxidized sample and 
brittle oxidized sample (500 mg/dm') was 
assessed as 100 J/cm 3 and 5 J/cnr, respectively 

4. Development of the 
E110 conservative 
oxidation kinetics 
performed in 

VNIINM, 1990 [42, 
43] 

The test data base obtained due to 
investigations stated in item 1 was used 

The following correlation was developed: 

( 10410^ 

A m = 920 exp-- . r > 

v T ) 

where Am-oxygen weight gain (mg/dm') 
T-temperature (K) 
x-time (s) 

5. High temperature 
oxidation tests of 
deformed claddings 
performed in 

VNIINM and VTE, 
1990-91 [42,43] 

The oxidation tests of pressurized (Ar) 
cladding samples at 700-850 C. The 
measurement of the cladding hoop 
strain in the ballooning area (the 
variation of the outer diameter relative 
increment was 0-85 % (AD/D 0 ) 

The systematic increase of the oxygen weight 
gain was observed as a function of AD/D 0 
increase. The dependence of the weight gain on 
the cladding surface in the ballooning area was 
estimated 

6. Thermal shock tests 
of oxidized claddings 
performed in 

VNIINM and VTE, 
1990-91 [42,43,44] 

The double-sided oxidation of El 10 
claddings at 800-1200 C and water 
quench cooling 

The fragmentation did not occur for all claddings 
with the ECR less than 18% (as-measured) 

7. Ring compression 
tests performed in 
VNIINM and VTE, 
1990 [42,43] 

The double-sided oxidation of El 10 
and Zry-4 (French and SANDVIK 
zircaloy) claddings at 800-1200 C and 
water quench cooling. (Reference tests 
with slowly cooled samples). 

Reference tests with the air oxidation 
of El 10 claddings. Ring compression 
tests of 30 mm oxidized claddings at 

20 C 

Ring compression tests showed that: 

• a sharp decrease of the El 10 ductility 
occurred after the achievement of some 
criterial ECR 

• this criterial ECR was a function of the 
oxidation temperature 

• the worst studied temperature was 1000 C 

• the best studied temperature was 800 C 

• the range of the El 10 zero ductility 
threshold may be estimated as 3^4% (as- 
measured) 


* All-Russian Research Institute of Inorganic Materials 
** All-Russian Heat Engineering Institute 

2.6 




















Test type 

Test characterization 

Test results 



• the Zry-4 oxidized cladding had the 
monotonic character of ductility reduction at 
the ECR increase. The zero ductility 
threshold of Zry-4 cladding was not higher 
than 15% ECR (as-measured) 



The visual observations of oxidized samples 
allowed to reveal that: 

• at ECRs close to the criterial value (S-4%) 
the white spalled oxide appeared on the 
E 110 cladding surface 

• Zry-4 oxidized samples were covered with 
the black bright oxide 

The comparative analysis of results of hydrogen 
content measurements showed that E 110 
claddings unlike Zry-4 had the tendency towards 
the high hydrogen absorption at 1000-1100 C. 
Reference tests with air oxidation confirmed that 
the E 110 residual ductility increased significantly 
in this case (without the cladding hydriding 
effect) 

The comparative data characterizing the 
sensitivity threshold to the cooling rate showed 
that this effect was insignificant 

The metallographic studies of the El 10 oxidized 

cladding allowed to note that: 

• the following difference in the El 10 and 
Zry-4 a-Zr(O) phase morphology was 
revealed: 

- the a-Zr(0) phase in the Zry-4 cladding 
consisted of equiaxed grains; 
the a-Zr(O) phase in the El 10 cladding 
consisted of thin plates 

• taking into account the difference in the 
phase transition temperatures for El 10 and 
Zry-4 alloys, the El 10 P-phase must dissolve 
more oxygen than the Zry-4 one to provide 
the a-Zr(O) phase initiation condition. This 
effect led to the difference in the a-Zr(O) 
thickness in El 10 and Zry-4 claddings 

8 . Thermal shock tests 
of VVER fuel rod 
simulators performed 
in VNIINM and 
RIAR\ 1998-2001 
[45, 46, 47, 48] 

The oxidation of unpressurized fuel 
rod simulators with the El 10 cladding 
at 900-1200 C, after that water quen¬ 
ching of oxidized simulators. Two 
types of simulators are used: 

• with AEO 3 fuel pellets and radiant 
heating of the fuel rod 

• with UO 2 fuel pellets and W-hea- 
ters installed inside the fuel stack 

One end of the fuel rod was open for 
the steam penetration 

Both types of VVER fuel rod simulators were not 
fragmented during and after thermal shock tests 
in the following range of test parameters: 

• 900-1200C 

• as-calculated ECR (using the E110 conserva¬ 
tive correlation) less or equal to 18% 

The different margin for the safety fragmentation 
threshold (18% ECR) was demonstrated as a 
function of the oxidation temperature and 
simulator type 


State Research Center "Research Institute of Atomic Reactors" 

2.7 















Test type 

Test characterization 

Test results 

9. Impact tests of 
oxidized VVER fuel 
rod simulators per¬ 
formed in VNIINM, 
1998-2001 [45-48] 

The E110 oxidized cladding after 
investigations performed as stated in 
the item 8 were tested in accordance 
with the following requirements 

The reference tests with the unoxidized cladding 
shown that the impact-toughness fracture was of 
64-89 J/cnr 


• claddings removed from the first 
type of fuel rod simulators were 
used 

• the circamferel notch 1-1.5 mm 
deep and 0.5 mm wide was made 
on each 100 mm oxidized cladding 

• the impact tests were performed at 
20 C 

Impact tests with the oxidized cladding 

demonstrated that: 

• the impact-toughness fracture was reduced 
down to 2.5 J/cnr at 5% ECR (as-calculated 
using the El 10 conservative correlation) 

• the impact-toughness fracture at 10-15% 
ECR (as-calculated) was about 1 J/cnr 

• the tough type of the cladding fracture was 
observed up to the 5% ECR 

• the brittle fracture occurred at 7% ECR 

10. Compression tests of 
E110 oxidized 
claddings performed 
in VNIINM, 1998- 
2001 [45-48] 

The double-sided oxidation of El 10 
claddings at 800-1200 C with water 
quenching. The compression tests of 
oxidized samples 30, 50 mm long at 
20-900 C 

Test data in general confirmed the results of 

previous tests: 

• the oxidation at 800 C was much better than 
the oxidation at 1000 C 

• the relative displacement at failure was 4% 
at the following combinations of as- 
calculated ECR and the oxidation 
temperature: 

- 800 C: 10% ECR 

1000 C: 5% ECR 

• no new effects were revealed in the tests of 
claddings taken from the fuel rod simulators 

The compression tests (at 20 C) of 
oxidized claddings (18% as-calculated 
ECR) taken from type two simulators 
tested in accordance with requirements 
given in the item 8 

The following effects of the mechanical test 
temperature were noted: 

• the ductility of El 10 oxidized at 200 C 
occurred if the ECR was less than 5% 

• at 18% ECR, the temperature effect in the 
range 20-500 C was insignificant 


It should be noted that intensive studies with the El 10 irradiated cladding were started in the middle of 1990s 
in addition to the test program presented in Table 2.2. The first results of investigations performed with the 
E110 irradiated claddings refabricated from VVER high bumup fuel rods (50 MW d/kg U) under LOCA 
conditions are presented in References 46^-8. 

The analysis of the whole scope of obtained results allowed to conclude that: 

1. Numerous thermal-shock tests performed with different El 10 samples (unirradiated and irradiated oxi¬ 
dized claddings and fuel rod simulators) showed that the El 10 fragmentation threshold was higher than 
18% ECR (calculated with the VNIINM conservative correlation) in the temperature range 800-1200 C. 

2. The effect of breakaway oxidation accompanied by the hydrogen uptake was revealed in El 10 claddings 
oxidized at temperatures higher than 800 C and relatively low ECRs. 

3. Mechanical tests (tensile, compression, impact) demonstrated that a sharp decrease in residual ductility 
of the El 10 oxidized cladding occurred in the measured ECR* range 4-7%. 


* The measured ECR should not be compared with the safety criterion (18%) because (as it was noted earlier) the calcu¬ 
lated ECR was used for the safety and analysis. That ECR was calculated using the conservative oxidation kinetics. 


2.8 











During 1990s several research programs devoted to the high temperature oxidation behavior of the El 10 
cladding were initiated in Germany, Hungary and the Czech Republic. Major findings of these test programs 
obtained by 2001 are presented in Table 2.3. 


Table 2.3. Major results of mechanical tests with the El 10 unirradiated oxidized claddings performed in 
_ Germany, Hungary • and Czech Republic during 1990-2000 


Test type 

Test characterization 

Test results 

1. Ring compression 
tests performed by 
J.Bohmert in Germa¬ 
ny (NC Rossendorf). 
1992 [49, 50] 

The double-sided oxidation of El 10 
and Zry-4 (SANDVIK) in water steam 
at 800-1100 C with the water quench 
cooling. Ring compression tests at 
20 C 

Results of mechanical tests showed that: 

• the ductility of the El 10 oxidized cladding 
(850-1100 C) decreased sharply down to the 
zero ductility threshold in 2(3)-4(5,6)% 
range of the as-measured ECR 

• the ductility of El 10 claddings oxidized at 
800 C was high up to the 8% ECR (the 
maximum value was achieved under this test 
conditions) 

• the Zry-4 ductility decreased monotonically 
as a function of ECR. relative displacement 
at failure was 8% at 18.5% ECR (as- 
measured) 

The visual observations showed that: 

• the white porous spalled oxide covered the 
E110 cladding sample at the relatively low 
ECRs 

• the black bright oxide covered the Zry 
samples 

The comparative measurements of hydrogen 
content in the oxidized claddings demonstrated 
that E110 had the tendency towards the high 
hydrogen absorption in contrast to the Zry-4 
cladding (especially at 1000 C). So. the 
maximum hydrogen content at the ECR of about 
18% (as-measured) was: 

• 2050 ppm in El 10 

• 130 ppm in Zry-4 

The microhardness measurements allowed to 
reveal that the microhardness (oxygen 
concentration) in the prior [5-phase was 
significantly higher in the El 10 cladding at the 
following test parameters: 

• temperature oxidation: 900-1100 C 

• oxidation time: 30 min 

2. Thermal shock tests 
performed in 

Hungary [51] 

The oxidation of 50 mm fuel rods 
simulators (El 10 cladding. AfCb 
pellet) at 1000-1250 C with the water 
quench cooling 

The cladding fragmentation did not occur if the 
ECR (calculated using the VNIINM conservative 
correlation) was less than 30% 

3. Ring compression 

tests performed in 
Hungary [51, 52] 

The double-sided oxidation of El 10 
and Zry-4 samples at 900-1200 C. 
Ring compression tests at 20 C 

Visual observations showed that: 

• the black bright oxide covered the Zry-4 
samples in most cases 

• the white spalled oxide covered El 10 
samples at relatively low ECRs 

The mechanical tests allowed to reveal that: 

• the Zry-4 ductility decreased monotonically 
at the ECR increase 

• the E110 ductility decreased sharply in the 
ECR range (as-measured) 1.6-5% 


2.9 














Test type 

Test characterization 

Test results 



The hydrogen content measurement showed that 
the Zry-4 hydrogen uptake was very low, the 
E110 hydrogen content achieved 600-800 ppm at 
the 5% ECR in the temperature range 900- 
1100 C. The E110 hydrogen absorption rate was 
somewhat slowed down at 1200 C 

4. Thermal shock tests 
performed in Czech 
Republic [53, 54] 

The double-sided oxidation of 30 mm 
E110 cladding samples at 800-1200 C 
with the water quench cooling 

The E110 cladding fragmentation threshold was 
observed at ECRs higher than 30% (as- 
measured) 

5. Ring compression 
tests performed in 
Czech Republic [53, 
54] 

The double-sided oxidation (argon and 
steam mixture) of 30 mm El 10 
cladding samples at 800-1200 C with 
the water quench cooling. Ring 
compression tests at 20 C 

The visual observations showed that: 

• the E110 steam oxidation at temperatures 
higher than 800 C led to the formation of the 
light color oxide and flaking-off effect (but 
without nodular corrosion effects) 

• the El 10 heating in the argon atmosphere up 
to the isothermal oxidation temperature led 
to the formation of the black lustrous oxide 
on the El 10 surface 

The ring compression tests and hydrogen content 

measurements allowed to reveal the following: 

• the zero ductility threshold of the El 10 
cladding was associated with 5% ECR (as- 
measured) 

• besides, the zero ductility condition was 
accompanied by the hydrogen content 500- 
700 ppm in the prior (3-phase 

• the further ECR increase led to the oxide 
spallation and the increase of hydrogen 
content up to 2000 ppm 


The experimental data organized in Table 2.3 confirmed that: 

• the tested El 10 claddings had a tendency towards breakaway oxidation and the hydrogen uptake at tem¬ 
peratures 850-1200 C; 

• a sharp decrease in El 10 ductility occurred at measured ECRs of 4-6%; 

• fragmentation of El 10 oxidized cladding was not observed at a calculated ECR less than 18% (1200 C) 
in accordance with the thermal-shock tests. 

Analysis of results of experimental investigations with the El 10 cladding performed in Russia and abroad 

lead to the following observations: 

1. All investigations presented in Table 2.2 and Table 2.3 were performed with as-received El 10 tubes 
manufactured during 1980s. But numerous improvements and changes were made during the 1990s 
in procedures for producing Zr ingots and fabricating El 10 cladding. 

2. The first published results characterizing the oxidation and mechanical behavior of other Zr-l%Nb 
cladding (French M5 alloy) showed that earlier breakaway oxidation and high hydrogen uptake were 
not observed in these tests with niobium-bearing cladding [55], 

3. Procedures of many oxidation tests (including the oxidation history, heating and cooling rates, tem¬ 
perature and weight gain measurements, steam flow rate value, etc.) were not validated and docu¬ 
mented along with the published data. The nature of cladding material used in investigations per¬ 
formed outside of Russia was not documented also. Procedures of mechanical tests, parameters of 
cladding samples and processing of obtained results were quite different in some cases and often un¬ 
known in other cases. 


2.10 










4. Taking into account that niobium-bearing alloys are of a high priority with respect to the increase of 
LWR fuel bumup, an understanding of physical phenomena defining the mechanical behavior of 
these claddings under accident conditions is an important international research task. 

These considerations have led to the decision to perform the special research program concerning the ex¬ 
perimental study of embrittlement ofZr-l%Nb (El 10) cladding under LOCA-relevant conditions. 


2.11 


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[50] Bohmert J., Dietrich M., Linek J. "Comparative Studies on High-Temperature Corrosion of Zr-l%Nb 
and Zircaloy-4", Nuclear Engineering and Design, 147 Nol, 1993. 

[51] Hozer Z., Maroti L., Matus L., Windberg P. "Experiments with VVER Fuels to Confirm Safety Crite¬ 
ria", Proc. of Top Fuel-2001 Meet., Stockholm, May 27-30, 2001. 

[52] Hozer Z., Griger A., Matius L., Vasaros L., Horvath M. "Effect of Hydrogen Content on the Embrittle¬ 
ment of ZR Alloys", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Transient 
and LOCA Conditions", Halden, Norway, September 10-14, 2001. 

[53] Vrtilkova V., Valach M., Molin L. "Oxiding and Hydrating Properties of Zr-l%Nb Cladding Material in 
Comparison with Zircaloys", Proc. of IAEA Technical Committee Meeting on "Influence of Water 
Chemistry> on Fuel Cladding Behavior", Rcz (Czech Republic), October 4-8, 1993. 


2.14 


[54] Vrtilkova V., Novotny L., Doucha R., Vesely J. “An Approach to the Alternative LOCA Embrittlement 
Criterion'’, Proc. of SEGFSM Topical Meeting on LOCA Fuel Issues, Argonne National Laboratory, 
May 2004 (NEA/CSNI/R(2004)19). 

[55] Brachet J., Pelchat J.. Hamon D.. Maury R., Jaques P.. Mardon J. "Mechanical Behavior at Room Tem¬ 
perature and Metallurgical Study of Low-Tin Zry-4 and M5™ (Zr-NbO) Alloys after Oxidation at 

1100°C and Quenching", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Tran¬ 
sient and LOCA Conditions ", Halden, Norway, September 10-14, 2001. 


2.15 

























































3. Oxidation behavior and embrittlement threshold of 

STANDARD E110 CLADDING: PROGRAM, TEST PROCEDURES, 

Discussion of test results 

3.1. The program concept and technical requirements to experimental works 


Previous investigations of the mechanical behavior of oxidized claddings under the LOCA conditions 
revealed a series of practical problems. Two of those are related directly to the present investigation: 

• a definite separation of such conceptions as the “cladding fragmentation threshold” and “cladding 
embrittlement threshold” is lacking on analyzing the study results; 

• the results of works performed by different researchers are discussed and compared without the analysis 
of test conditions and regard for peculiar features of procedures employed to determine the output 
parameters. 

As for the first of the problems, the following approach was proposed to specify two criterion conditions of 
the cladding: 

1. The cladding fragmentation threshold characterizes the actual loss of the cladding structural integrity 
under the impact of a complex combination of loading factors. This threshold can be determined for 
each of the typical sections of the accident scenario and for each of the post-accident actions only 
with the help of well-planned integral experiments with fuel rods and simulators of fuel rods. 

2. The embrittlement threshold characterizes the conditions under which the cladding ductility is close 
to zero. In this case, it is of importance that the appropriate threshold value be determined by 
relatively simple parametric experiments including those using uniaxial mechanical tests. 

It is evident that, from the theoretical point of view, the cladding fragmentation threshold should be the 
major subject of investigation within the context of this program. However, in spite of the fact that almost 30 
years after the time when the first licensing criterion for safety was developed, the following basic 
conclusion about the basis of the criterion has not changed up to now: “There is some lack of certainty as to 
just what nature of stresses would be encountered during the LOCA” [1]. In turn, the insufficiency of such 
knowledge leads to the difficulty of justifying requirements on the basis of integral experiments. Therefore, 
with due respect to the importance of such experiments as a thermal shock, it should be acknowledged that 
the experimental determination of the embrittlement threshold for claddings -- with the subsequent 
introduction of coefficients regulating the necessary minimum margin for the residual ductility in oxidized 
claddings -- was and remains the basic method for the practical safety studies. 

Therefore, the further analysis of possible variants for the experiments was restricted to the following list of 
mechanical tests for oxidized claddings: 

• impact tests; 

• three-point bending tests; 

• tensile tests; 

• ring compression tests. 

The consideration of these variants showed that: 

• impact tests do not allow one to obtain a stress-strain curve in the explicit form and. besides, these tests 
are not standardized so that one could not directly compare the results obtained for different claddings in 
different countries; 

• three-point bending tests present a very promising type of study; however, the procedure requirements 
are not quite developed and the comparative data base for these test is practically lacking; 


3.1 


• tensile tests present the most valuable type of mechanical tests; however, tensile tests of claddings with a 
high level of oxidation require to improve of test techniques and, besides, a comparative data base 
obtained from these tests is extremely limited; 

• ring compression tests present the simplest and accessible type of tests allowing to obtain good statistical 
data within a limited time. Therefore, this type of tests has a representative data base containing the 
results of previous investigations of alloys El 10, Zry-4, M5, Zirlo. 

Taking into account all stated above, the final choice of the method to determine the embrittlement threshold 
for the El 10 cladding was made in favor of ring compression tests. Nevertheless, it should be noted that this 
type of tests has a series of disadvantages the basic ones of which are as follows: 

• the lack of standard requirements imposed on geometrical sizes of tested specimens (specimen length); 

• the lack of physically justified rules for the interpretation of test results. 

Therefore, special scoping tests were stipulated in the program of these studies. The aim of these tests was to 
obtain a data base necessary to specify the requirements to the test conditions and to justify the processing 
procedures for obtained results. It should be noted that the list of procedural issues revealed at the stage of 
this program development included several problems associated with the oxidation conditions. 

The first of these problems may be characterized by the following propositions: 

• from the formal point of view, the El 10 alloy embrittlement threshold must be determined and compared 
with the license embrittlement criterion; 

• the license embrittlement criterion for the El 10 alloy (18%) is based on the employment of the VNIINM 
conservative correlation for the El 10 oxidation kinetic description [2]; 

• just the same approach is used for Zircaloy-4 alloy (15% in France and 17% in U.S.) but in this case the 
Baker-Just conservative correlation is often employed [3]; 

• both conservative correlations were developed according to individual (as applied to each alloy) rules, 
and the conservatism degree was not necessarily the same. Therefore, a direct comparison of the 
mechanical behavior of these two alloys (and others such as M5 and Zirlo) with one of the conservative 
correlations is not meaningful from the physics point of view; 

• it is evident that the practical solution of this problem may be provided on moving from the conservative 
correlations to best-estimated correlations with a simultaneous transformation of the license criteria into 
the same system of coordinates (i.e., as-measured); 

• in this case, the proposed approach allows to solve one more important problem, namely, to obtain a data 
base characterizing the behavior of claddings fabricated from different alloys in the system of 
coordinates (ECR vs. time), as it is evident that both these factors are competitively significant in the 
consideration of accident processes. 

However, earlier work demonstrated that different researchers developed a number of different correlations 
for the El 10 alloy oxidation kinetics [4, 5, 6, 7, 8, 9, 10]. Therefore, it was decided to obtain additional data 
that would allow us to pick the best estimated correlation from those. The initial program stage would thus 
include the following: 

• provide a sufficient quantity of oxidation tests in the research Program necessary to obtain additional 
data on the E-110 alloy oxidation kinetics; 

• perform oxidation within the range of 5-20% ECR. 

The final stage of the scoping tests was dedicated to determination of requirements to be imposed on the 
types of claddings to be tested, on the oxidation conditions and on the parameters of mechanical tests with 
regard to such potentially significant factors as: 

• oxidation type (single-sided or double-sided); 

• oxidation temperature; 

• heating and cooling rates; 


3.2 


• cladding material (as-received El 10. El 10K. E635 tubes, as-received El 10 cladding (El 10A), irradiated 
commercial El 10 cladding); 

• temperature of mechanical tests. 

To optimize the quantity of tests, it was decided to divide the Program into three relatively independent 
stages: 

1. Scoping tests. 

2. Reference tests. 

3. Sensitivity studies. 

The goals, tasks and results of scoping tests will be discussed in the section 3.2 of the report. Table 3.1 
presents the logical principles laid into the basis of the development of reference tests and sensitivity studies 
as well as the parameters of these tests. 


Table 3.1. The list of tasks and technical requirements 


Program stage 

Major tasks 

Technical requirements 

Motivation 

1. Reference 

tests 

To determine the zero ductility 
threshold of El 10 cladding as a 
function of the following 
oxidation parameters: 

• cladding heating rate 

• cladding cooling rate 

• ECR 

To find the optimal test mode for 
the sensitivity studies 

1.1. Varied parameters: 

• heating rate: 

0.5 C/s 

- 25 C/s 

• cooling rate: 

0.5 C/s 

- 25 C/s 

- 200(170-270) C/s 

- ECR: 

1.2. Fixed parameters: 

• cladding type: as-received 
E110 cladding tube 

• coolant (oxidation medi¬ 
um): water steam (0.1 MPa) 

• oxidation type: double¬ 
sided 

• hold temperature: 1100 C 

• combination of oxidation 
modes: 

- F/F (fast heating (25 C/s) 
and fast cooling (25 C/s)) 

- F/Q (fast heating (25 C/s) 

and quench cooling 

(200 C/s)) 

- S/S (slow heating (0.5 C/s) 
and slow cooling (0.5 C/s)) 

- F/S (fast heating (25 C/s) 
and slow cooling (0.5 C/s)) 

- S/F (slow heating (0.5 C/s) 
and fast cooling (25 C/s)) 

• parameters of mechanical 
tests: 

- 20 C 

1 mm/min 

• Different heating 
rates allow to 
determine the 
sensitivity of 
embrittlement 
phenomena to the 
initial stage of 
cladding oxidation 
history (oxidation in 
the a-(3 temperature 
range) 

• Different cooling rates 

allow to determine the 
sensitivity of 

embrittlement 
phenomena to: 

- (3—>a' phase transient 
conditions 

- thermal stresses under 
cooling conditions 

• Five combinations of 
heating and cooling 
rates (F/E, F/Q. S/S, 
F/S, S/F) allow to re¬ 
veal definitely the im¬ 
pact of appropriate pa¬ 
rameters on mechanical 
properties of claddings 


3.3 







Program stage 

Major tasks 

Technical requirements 

Motivation 

2. Sensitivity 
studies 

2.1. Tests with 
Zry-4 cladding 

To perform several oxidation and 
mechanical tests with the Zry-4 
cladding and to compare the 
obtained results with the 
published Zry-4 data 

2.1.1. Varied parameters: 

• combination of heating and 
cooling rates: 

- S/S 

- F/F 

• ECR: 11-12% 

2.1.2. Fixed parameters: 

• cladding type: as-received 
Zry-4 

• coolant: water steam 

• oxidation type: double¬ 
sided 

• hold temperature: 1100 C 

• parameters of mechanical 
tests: 

- 20 C 

- 1 mm/min 

• Comparison of Zry-4 

data obtained in the 
frame of this work with 
the Zry-4 data obtained 
previously allows to 

verify the whole set of 
experimental 
procedures 

• Comparison of the 

E110 and Zry-4 data 
allows to reveal general 
differences in the 

physical behavior of 
these alloys under 
oxidation conditions 

• Two combinations of 
heating and cooling 
rates allow to reveal the 
sensitivity of the Zry-4 
cladding to oxidation 
conditions 

2.2. Oxidation 
type 

To perform a single-sided 
oxidation of El 10 cladding and to 
compare single-sided and double¬ 
sided test results 

2.2.1. Varied parameters: 

• ECR range: 7-12 % 

2.2.2. Fixed parameters: 

• cladding type: as-received 
E110 tube 

• coolant: water steam 

• oxidation type: single-sided 

• hold temperature: 1100 C 

• F/F combination of heating 
and cooling rates 

• parameters of mechanical 
tests: 

- 20 C 

- 1 mm/min 

• A single-sided oxida¬ 
tion characterizes 

cladding behavior in 
the undeformed part of 
the rod (away from 
balloon) 

• French data base on the 
embrittlement threshold 
of the M5 alloy has 
been developed for a 
single-sided oxidation 

• Comparative data base 

(single-sided, double¬ 
sided oxidized 

claddings) allows to 
estimate the 

dependence of test 
results on the oxidation 
type 

2.3. Oxidation 
temperature 

To perform a double-sided 
oxidation of El 10 cladding at 
different hold temperatures and to 
reveal the sensitivity of the 
cladding behavior to this 
parameter 

2.3.1. Varied parameters: 

• hold temperature: 

- 800,900,950,1000,1100, 
1200 C 

• ECR: 6-13 % 

2.3.2. Fixed parameters: 

• cladding type: as-received 
E110 tube 

• coolant: water steam 

• oxidation type: double¬ 
sided 

• F/F combination of heating 
and cooling rates 

• parameters of mechanical 
tests: 

- 20 C 

1 mm/min 

• The El 10 
embrittlement 
phenomena could be 
sensitive to the hold 
temperature in 
accordance with the 
previous test data base 

• Comparative data base 
obtained for different 
temperatures allows to 
estimate this effect 


3.4 









Program stage 

Major tasks 

Technical requirements 

Motivation 

2.4. Alloying 
composition 

To perform a double-sided 
oxidation of as-received E110K. 
E635 cladding tubes and to 
determine the sensitivity of 
cladding behavior to alloying 
composition 

2.4.1. Varied parameters: 

• cladding material type: 

- E110K 

- E635 

• ECR: 7-14 % 

2.4.2. Fixed parameters: 

• coolant: water steam 

• oxidation type: double¬ 
sided 

• F/F combination of heating 
and cooling rates 

• hold temperature: 1100 C 

• parameters of mechanical 
tests: 

- 20 C 

- 1 mmy min 

• The E11 OK alloy has a 
high concentration of 
oxygen (approximately 
the same as the M5 
alloy has). Oxygen is 
considered as an 
alloying component in 
case of concentrations 
like that 

• The E635 alloy has Fe 
and Sn as alloying 
components (similar to 
those in the Zirlo alloy) 

• The comparative data 
base with El 10, 

E110K, E635 test 
results allows to 
understand the 
sensitivity of oxidation 
behavior of Zr-Nb 
alloys to the major 
chemical components 
of the cladding material 

2.5. Tempera¬ 
ture of 

mechanical 

tests 

To perform ring compression tests 
at 135 C and several ring tensile 
tests at different temperatures and 
to determine the sensitivity of 
embrittlement phenomena to the 
temperature of mechanical tests 

2.5.1. Varied parameters: 

• temperature of ring 

compression tests: 20. 135, 
200, 300 C 

• temperature of ring tensile 
tests: 135, 200, 300 C 

2.5.2. Fixed parameters: 

• E110 oxidized cladding 
tubes after a double-sided 
oxidation at 1100 C and F/F 
combination of heating and 
cooling rates 

• The first safety 

embrittlement criterion 
(17% ECR) was 
validated at 135 C (the 
saturation temperature 
during the reflood) 

• Possible earlier 

embrittlement of the 
E110 alloy may be 
associated with the 
generation of hydrides 
at low temperatures. 
But it is known that 
hydrogen solubility in 
zirconium alloys is a 
strict function of the 
temperature 

• The comparison of 

mechanical tests 

performed at different 
temperatures allows to 
reveal the sensitivity of 
the E110 embrittlement 
behavior to post-LOCA 
conditions 


3.5 








Program stage 

Major tasks 

Technical requirements 

Motivation 

2.6. Type of 

mechanical 

tests 

To perform several ring tensile 
tests, several three-point bending 
tests and to develop the data base 
characterizing the representativity 
of the E110 embrittlement 
threshold determined due to ring 
compression tests 

2.6.1. Varied parameters: 

• type of mechanical tests: 

- ring tensile tests; 

- three-point bending tests 

2.6.2. Fixed parameters: 

• the El 10 oxidized cladding 

after a double-sided 

oxidation at 1100 C and F/F 
combination of heating and 
cooling rates 

• The comparative data 
obtained for three types 
of mechanical tests 
(compression, tensile, 
bending) allow to esti¬ 
mate the representativi¬ 
ty of oxidation limits 
developed on the basis 
of ring compression 
tests. 

2.7. Fuel 
bumup 

To perform the scoping tests 
(oxidation and ring compression) 
with E110 commercial irradiated 
claddings refabricated from 

VVER fuel rods and to compare 
obtained results with the data base 
on as-received claddings 

2.7.1. Varied parameters: 

• hold temperature: 1000, 
1100,1200 C 

• combination of heating and 
cooling rates: S/S, S/F, F/F 

• ECR: 6-16% (as- 

measured) 

2.7.2. Fixed parameters: 

• cladding type: irradiated 
E110 cladding made from 
commercial fuel rods with 
bumup ~50 MWd/kg U 

• coolant (oxidation 

medium): water steam 

• oxidation type: double¬ 
sided 

• parameters of mechanical 
tests: 

- 20, 135 C 

1 mm/m in 

• The preliminary data 
base obtained in these 
tests will allow to 
estimate high bumup 
fuel effects in the 
context of 

embrittlement pheno¬ 
mena of the E110 
cladding 


3.2. Methodological aspects of oxidation and mechanical tests 


3.2.1. Oxidation tests 

In accordance with the review results discussed in Chapter 2 oxidation tests have a long historical tradition 

based on the following approach: 

• oxidation is performed under indirect heating of a cladding sample; 

• cladding samples are oxidized and the same samples are used in the mechanical tests; 

• oxygen weight gain is determined by weighing the sample during the oxidation process or by weighing 
the sample before and after the test. 

However, the analysis performed on preparing this paper has shown that this approach has several 

disadvantages, the following two of which are the basic ones: 

1. Oxidation of the cladding samples of any length is accompanied by the occurrence of end-effects caused 
by oxygen (hydrogen) absorption at the end-faces of the cladding sample. These end-effects lead to two 
consequences of importance: 


3.6 








1.1. These parts ot the cladding sample are characterized by a higher tendency to the embrittlement 
caused by: a) a higher ECR and b) additional stresses in the Zr0 2 layer resulting in an early 
initiation of the breakaway oxidation: 

1.2. The ECR (weight gain) of the oxidized cladding sample is overestimated as the area of the cladding 
end-taces is not taken into account in the ECR (weight gain) calculations. 

It is evident that the results of oxidation and mechanical tests are the more sensitive to the end-effects, 
the less is the length of the cladding samples. This approach is especially unacceptable for El 10 
claddings as the end-effects intensify the development of the breakaw ay phenomena. 

Besides, the employment of short cladding samples extends the duration and labour-intensiveness of the 
investigation, on the one hand, and restricts the opportunity to perform other types of tests (except for 
mechanical ones) with oxidized samples, on the other hand. 

2. Determination of the weight gain by measuring the difference in the sample weight before and after 
oxidation is quite an undesirable procedure for alloys susceptible to such effect as spallation of ZrO:. 
This method application for such alloys leads to the underestimation of the actual weight gain. 

Basing on this analysis results, on technical requirements presented in Table 3.1. the following major 
methodical problems have been formulated: 

1. To develop the facility and procedure necessary for oxidation of the cladding sample 100 mm long; 

2. To develop the experimental procedure independent of end-effects and ZrO : spallation and flaking off to 
determine the ECR (weight gain). 

A brief description of this work results is presented in the following sections. 


3.2.1.1. Test samples and oxidation apparatus 


Fig. 3.1 illustrates the appearance of test samples developed for this program. 


Window for steam outlet 


Sample operating length (100 mm) 


Water 
steam 

Sample attachment point to pull rod 


Fig. 3.1. Schematics of the cladding sample 



The sample shown in Figure is intended for the double-sided oxidation. The sample for single-sided 
oxidation had no window' for steam outlet and. besides, the sample bottom end was plugged. The sample was 
oxidized in the facility the basic diagram of w hich is demonstrated in Fig. 3.2. 


3.7 

















f-besl-t.cdr 


c =L 


Fig. 3.2. Oxidation test apparatus 


Description of the facility and procedures employed for the oxidation of samples are described in more detail 
in Appendixes A-l, A-2. Parameters of all types of the cladding samples used in this work are presented in 
Appendix A-3. Results of scoping tests performed to characterize the temperature distribution in the 
cladding sample are given in Appendix A^4. In accordance with the Program of research, the oxidation 
facility provided for the following combinations of heating and cooling rates: S/S, S/F, F/S, F/F, F/Q, where 
S is slow, F is fast, and Q is quench. Fig. 3.3 demonstrates typical examples for the temperature histories for 
these combinations, p.3.8. It should be noted also, that the preliminary heating of cladding samples to 150 C 
was performed with the argon flow, and the further heating was performed with the water steam flow. 


3.8 



















































































































Combinations of oxidation modes 


F/F 



1400 


1200 

u 

1000 

•— 

800 

ca 

i_ 

u 

Q. 

*— 

600 

tJ 

400 


200 


















/ 















J 










#41 

ECR=8.6 





%r 

























50 100 150 200 250 300 350 400 

Time (s) 


F/S 


1400 

1200 

1000 


£ 800 


Sr 600 


u 

f— 


400 

200 

0 






#40 

ECR 

=8.7 % 

- 






























— 

— 

— 




200 400 600 800 1000 1200 

Time (s) 


F/Q 


S/F 



1400 


1200 


1000 

G 


p 

800 

s 


6 

c, 

»— 

600 

S 

f- 

400 


200 





















r 







i 

#71 






£.L.K =: V./ /o 











peek 7 1 -v2-en.gr/ 







1400 

1200 

1000 


s 800 


'-I 

C- 

E 

« 

H 


600 

400 

200 


















1 







1 







T 









\ 



#34 

ECR=7. 

2% 







A 














100 200 300 400 

Time (s) 


500 


600 


0 

600 800 1000 


1200 1400 

Time (s) 


1600 1800 2000 


S/S 



1600 


U 

'd 

1200 

5 

o 

Q. 

r— 

800 

H 

D 

f— 

400 























#38 









rr 

c. 

=7.6 % 






400 


800 1200 1600 2000 

Time (s) 


2400 


2800 


3200 


3600 


Fig. 3.3. Types of temperature histories for different combinations of heating and cooling rates 


Appendix A-5 contains the description of the method for the ECR determination. Major provisions of the 
method developed in the frame of this work may be characterized by the following way: 

AW 0 =AW,-AW e , 

where AW 0 - the cladding oxygen weight gain obtained during the oxidation test (mg/cm 2 ); 

AW-, - the oxygen weight gain calculated for the case with the cladding oxidation during the 
infinite time (full transformation of Zr-l%Nb to Zr0 2 and Nb 2 0 5 oxides) (mg/cm 2 ); 

AW e - the oxygen weight gain necessary to oxidize the metallic part of the cladding remaining 


3.9 




























































































after the oxidation test to the stoichiometric oxide. 

The concept of this method consists in the fact that a new experimental procedure named the “extra 
oxidation” is introduced in addition to the oxidation test. This procedure includes the following stages: 

• a ring sample (with the measured length) cut out from the oxidized cladding sample of the initial length 
100 mm is located into the crucible of the oxidation facility; 

• the crucible with the ring sample is weighed; 

• the extra oxidation of the ring sample is carried out up to the complete oxidation of a metallic phase and 
the generation of a stoichiometric oxide; 

• the crucible with the extra oxidized ring sample is weighed and oxygen weight gain obtained during the 
extra oxidation (AW e ) is determined. 

Thus, the oxygen weight gain during the oxidation test is determined as the difference between the 
theoretical value of weight gain necessary for the complete oxidation of Zr-l%Nb ring sample (AWoo), and 
the measured weight gain after the extra oxidation procedure (AW e ). The introduction of this procedure for 
the weight gain determination and employment of a long (100 mm) cladding sample for the oxidation tests 
allowed to solve the following urgent investigation problems: 

• to fabricate several ring samples for mechanical tests from one oxidized sample; 

• to perform mechanical tests of ring samples at different temperatures; 

• to fabricate special samples from the same oxidized sample for metallographic investigations, 
fractography research, hydrogen content measurement, scanning electron microscopy examination 
(SEM), transmission electron microscopy examination (TEM). 

Fig. 3.4 presents the major provisions for experimental procedures with cladding samples. 


Manufacturing of cladding sample (100 mm) and clearing of sample surface 

The cladding sample oxidation - 


XX 


Cutting of the oxidized sample and separation of ring samples for different investigations 


Ring samples for 
mechanical 
tests: 

1 2 ... / 

-U- 

Mechanical 

tests 

Extra oxidation 
of samples 
and ECR 
determination 


Ring samples for 
metallography 
examination: 
12 ... k 




W 

..rHp^ 

* 

I / 

Umm/ 


$ 


i 


_-=XJ=’~ 

Ring samples for 
hydrogen content 
measurement: 
l 2 ... n 


Appropriate investigations 


Test data base 


Ring samples for 
other types 
of investigations 



Fig. 3.4. Development of the data base with test results 


The procedure developed for mechanical tests is described in the following paragraph. Other special 
procedures are characterized in Appendix A-6. 


3.10 














3.2.2. Mechanical tests 


In accordance with the program, three types of mechanical tests were provided to assess the mechanical 
behavior of the El 10 oxidized cladding: 

1. Ring compression tests. 

2. Ring tensile tests. 

3. Three-point bending tests. 

In this case, ring compression tests were considered as the basic type of mechanical tests. And two other 
types of tests were intended to assess the representativeness of the obtained data base. 


3.2.2.1. Ring compression tests 

The procedure ot ring compression tests may be schematically described in the following way (see Fig. 3.5): 

• an oxidized ring sample is located in the test machine; 

• a ring sample is compressed at the constant rate of cross-head displacement; 

• the load-displacement diagram is noted during the test. 

Compression 
load 

T 


Fig. 3.5. Test machine for ring compression tests of oxidized cladding samples 

As it was already mentioned above, this method is not used (as a rule) for the measurement of mechanical 
properties. How ever, it allows to determine that the threshold state of oxidized metal at which the cladding 
failure takes place without macroscopically significant plastic strain. This cladding material may be 
considered brittle. Accordingly, the ECR of this material is regarded as the cladding embrittlement threshold. 

The analysis of previous approaches used to determine the embrittlement threshold of zirconium claddings 
on the basis of ring compression tests (see Chapter 2) allowed to reveal the following general problems: 

• cladding samples of different length (6-30 mm) were used for these tests; 

• the standard procedure for the processing of load-displacement diagrams was not developed; 

• the cladding embrittlement threshold was estimated by different methods basing on results of ring 
compression tests; 

• the organization of previous data into the summarized data base is not possible for any tested alloys (see 
the additional information presented in section 3.3.2). 



3.11 










Therefore, thorough development and validation of the procedure for ring compression tests were the most 
important tasks in this work. 

A real load-displacement diagram based on the ring compression test of the El 10 oxidized cladding allows to 
illustrate the above mentioned problems more obviously (see Fig. 3.6) 



Fig. 3.6. The as-measured load-displacement diagram of the El 10 oxidized cladding sample 

The list of three questions formulated in Fig. 3.6 should be added with the fourth one: what is the impact of 
the ring sample length on the parameters of the load-displacement diagram? To find answers to these 
questions, the program of scoping mechanical tests was worked out the major tasks of which were as follows: 

• the determination of an effective elasticity modulus for the ring sample; 

• the development of the processing procedure for as-measured load-displacement diagrams; 

• the sensitivity studies to validate a standard length of the ring sample; 

• the development of a special data base necessary to determine the relationship between the cladding 
fracture and load-displacement response. 

The first step of this work was devoted to the transformation of the as-measured load-grip displacement 
diagram into the standard form using (for each tested ring sample) the procedure of the zero-displacement 
point determination in accordance with the approach presented in Fig. 3.7. 


3.12 































Fig. 3.7. Procedure for the determination of the sample zero-displacement point 


After that, special scoping tests were performed to determine the verification data characterizing Young’s 
modulus of the cladding ring sample. It should be noted that the compression diagram consists of two 
portions: 

1. The portion of the sample elastic deformation with the linear relationship between the load and dis¬ 
placement. 

2. Nonlinear portion characterizing the plastic component accumulated in the total deformation of sample. 
The effective modulus of elasticity is employed in the uniaxial mechanical tests to determine the plastic 
component of a total elongation. 

However it is known that the ring sample compression induces not only the compression and tensile stresses 
but also bending stresses that are distributed along the sample perimeter in the compound way. In this case, it 
should be pointed out that the investigation of the similar issue performed earlier allowed to develop the 
procedure of comparison of the effective elastic modulus and the load relief lines [11], Fig. 3.8 illustrates the 
results of appropriate scoping compression tests. 


3.13 













Displacement (mm) 


Fig. 3.8. Determination of effective modulus of elasticity basing on the results 

of scoping compression tests 


The procedure of this test included the compression of the ring sample with the load reduction in several 
points of the diagram (partial load relief)- It is known that the process of the load reduction in the elastic 
deformation area is described by the similar straight line as that describing the sample compression process. 
The comparative analysis of the obtained data has shown that the linear portion of the load-displacement 
diagram characterizes the real elastic deformation of the sample (tgai=stga 2 . ..i). Thus, this part of the diagram 
can be used to determine the plastic component of the deformation on the diagram processing. 

The next step of the procedure development was devoted to the determination of the parameter characterizing 
the margin of residual ductility in the oxidized cladding. Fig. 3.9 presents the major provisions of the 
procedure proposed for this work. 

In accordance with this procedure, the cladding strain at the compression was assessed by two parameters: 

1. Relative displacement at failure (S t ). 

2. Plastic component of relative displacement at failure (S p i). 


3.14 





















A/- grip displacement (mm) 

5- relative displacement (A l/D a 100 %) 

Sf- relative displacement at failure 

S p i - plastic component of relative displacement 

Fig. 3.9. The processing procedure for the load-displacement diagram of the ring compression test 

of the E110 oxidized cladding 


The first of these parameters was used in some of the early ring compression tests [2, 12, 13, 14, 15, 16, 17, 

18]. However this approach has several essential weak points: 

• the sensitivity of this parameters to the deformation plastic component is reduced as the cladding 
embrittlement is progressing; 

• this parameter does not tend to zero as the cladding approaches the zero-ductility threshold; 

• it is impossible to determine the zero-ductility threshold on the basis of the cladding relative 
displacement as a function of the ECR without an additional relationship connecting these two 
parameters. 

Data presented in Fig. 3.10 illustrate these problems obviously. 



Fig. 3.10. Relative displacement at failure of the El 10 cladding after a double-sided oxidation and S/S 

combination of heating and cooling rates as a function of the ECR 


3.15 
































The analysis performed by the report authors to solve these problems has shown that from the physics point 
of view the macroscopic zero-ductility threshold or, in other words, the macroscopic cladding embrittlement 
threshold may be defined as such ECR critical value at which the plastic component of the cladding strain at 
ring compression tests tends to zero. Fig. 3.11 presents the graphic interpretation of this approach for the 
same data base demonstrated in Fig. 3.10. 



40 


30 


20 


10 


0 





s 

s' 

z 

js 

S’ 

s 



— 

/ 

/ 

/ 

s 

S' 

>* 



s 

z 

s' 

S' 

S 



/ The zero-ductility threshold using 

. the consen’ative approach 



10 


20 


30 


40 


Relative displacement at failure (%) 

Fig. 3.11. The validation of the procedure for the zero-ductility threshold determination 


It should be noted that the term “the plastic component of relative displacement at failure” reflects clearly the 
physics sense of the appropriate parameter, however, the understanding of this term requires the knowledge 
of the ring compression test procedure in detail. Therefore, to provide for the understanding of the 
investigation results by a wide circle of specialists interested in this work, it was decided to replace this term 
by the term “residual ductility”. The term “residual ductility” adequately reproduces the sense of the subject 
of investigation, on the one hand, and does not require a special study of the procedure for mechanical tests, 
on the other hand. 

The next task of scoping mechanical tests was devoted to the development of the comparative data base for 
the assessment of sensitivity of the test results to the sample length. The appropriate tests were performed 
with ring samples the length of which was varied from 8 up to 25 mm. Two oxidized samples 100 mm long 
were selected for these tests: 

• the first sample was characterized by a high margin of residual ductility; 

• the second sample was practically brittle. 

Fig. 3.12, Fig. 3.13 show the results of these scoping tests. 


3.16 














Fig. 3.12. Sensitivity of a relative displacement at failure to the ring sample length 

for the ductile sample 



Fig. 3.13. Sensitivity of a relative displacement at failure to the ring sample length 

for the brittle sample 

The analysis of obtained data has allowed to make the following practical conclusion: the relative 
displacement at failure is independent of the ring sample length. In this case, the increase of the effective 
modulus of elasticity and maximum load with the sample elongation is quite natural because this effect is 
explained by a higher stiffness of longer samples. However, it should be pointed out that this conclusion 
might not be applied to the case when samples of different length were oxidized. As it was already stated, the 
mechanical behavior of these samples is complicated by the occurrence of end effects (if end parts of 
oxidized sample were not cut). 

The final task of this stage of work was formulated as follows: to develop the experimental data necessary to 
identify the cladding fracture in load-displacement diagrams obtained from ring compression tests. The 
appropriate analysis showed that in spite of the evident importance of this issue, the purposeful studies of this 
subject were performed in neither of the previous programs. The main point of the problem under discussion 
may be illustrated by the examples presented in Fig. 3.6 and Fig. 3.14. 


3.17 











































































The obtained results may be divided into two groups: 

1. Diagrams with relatively smooth load variation (two upper diagrams in Fig. 3.14). 

2. Diagrams with a sharp drop (20% and more) of load with 2-15 % relative displacement (two lower 
diagrams in Fig. 3.14 and diagram in Fig. 3.6). 

Several special tests were performed to interpret sufficiently diagrams of different types. Metallographic 
cross-sections were prepared from the samples of high residual ductility (a smooth diagram load- 
displacement). Besides, reference oxidation tests were performed with the ECR two characteristic levels: 

3. ECR=5.8 %. The sample with this ECR has a high margin of ductility. 

4. ECR=9.9 %. The sample with this ECR is brittle. 

Several ring samples were fabricated from each oxidized sample 100 mm long. After that, ring compression 
tests were performed in accordance with the following outline: 

• the ring sample loading up to the occurrence of the first drop of load in the diagram and termination of 
this sample test; 

• the test of the next sample from this batch up to the occurrence of the second drop of load in the diagram 
and termination of the test; 

• and so on. 

This approach allowed to visualize the state of the cladding sample in all characteristic points of load- 
displacement diagrams. Fig. 3.15 demonstrates a typical appearance and cross-section of tested ring samples 
with a high ductility margin. 


3.18 
























































1200 




a) unoxidized cladding (appearance of a ring sample after mechanical tests) 



View A 




Relative displacement (%) 

b) oxidized cladding (cross-section of a ring sample after mechanical tests) 

Fig. 3.15. Demonstration of the state of ductile claddings after ring compression tests 


In accordance with the obtained data, the compression of the unoxidized cladding sample results in its plastic 
deformation with no indications of failure. The shape of deformed oxidized sample with a high margin of 
ductility is practically the same as that of the unoxidized sample. However, the comparative analysis of such 
two phenomena as the drop of load after approximately 40 % displacement and the cross-section view allows 
to state that two nonthrough sample ruptures appeared in the areas of the maximum tensile stresses. Thus, the 
low in value and smooth decrease of load indicates the occurrence of a local rupture in the cladding outer 
layer prior to the P phase. The analysis of the oxidized cladding diagram presented in Fig. 3.15 allows to 
formulate one more question: what means the local drop of load noted in the diagram on the transfer from the 
elastic deformation to the plastic deformation (see view A). This question is of sufficient importance because 
this characteristic area was present in the majority of the load-displacement diagrams. 

To answer this question, a series of tests was performed with two rings fabricated from the same sample as 
the ring shown in Fig. 3.15. Fig. 3.16 presents the results of these tests. The analysis of obtained data shows 
that the observed load peak is not associated with the cladding failure. This peak indicates the process of 
formation of microcracks on the cladding surface in the areas of concentrated tensile stresses. The length of 
these cracks corresponds to ZrO : and a-Zr(O) layer thickness. Cross-sections of the ring compressed up to 
the maximum load (Fig. 3.16b) demonstrate that a growth and some opening of these cracks take place, 
however, the cladding metal part (the prior P-phase) continues to be deformed plastically. 

The next stage of this study was dedicated to the interpretation of load-displacement diagrams having 
segments with a sharp drop of load. The first type of these diagrams reflects the behavior of oxidized samples 
with a definite ductility margin (see Fig. 3.6. Fig. 3.14 The lower left diagram). The second type of diagrams 
characterizes the behavior of a brittle cladding (Fig. 3.14 The lower right diagram). In accordance with this 
classification, the failure of the following two samples was investigated: 

• Zry-4 sample #43; 

• El 10 sample #25. 


3.19 




















Relative displacement (%) 


Relative displacement (%) 



Fig. 3.16. The data base for the interpretation of load-displacement diagrams for oxidized samples 

with a high ductility margin 


The load displacement diagram of Zry-4 samples oxidized up to 11.3 % ECR at S/S combination of heating 
and cooling rates is shown in Fig. 3.17. 


3.20 











































































800 


600- 


■g 400 
o 


200 - 


0 


#43-2 


4.6 


5.0 


General question: 

When <>cl nr with cladding at this point of diagram? 


5.4 


7.4 


7.8 


5.8 6.2 6.6 7.0 

As-measured displacement (mm) 

Fig. 3.17. The load-displacement diagram of Zry-4 cladding sample 

with a partial ductility’ margin 


8.2 


To get the answer to the question put in Fig. 3.17, additional mechanical tests of two ring samples cut out of 
the oxidized sample were performed. Fig. 3.18, Fig. 3.19 demonstrate the results of these tests. 





Fig. 3.18. The data base to characterize the mechanical behavior of the cladding sample 

with a partial residual ductility before the fracture 


3.21 

















Fig. 3.19. The data base to characterize the mechanical behavior of cladding sample with the partial 

residual ductility at failure 


The compression of the ring sample tested in the first case (Fig. 3.18) was terminated immediately after the 
sharp load drop was noted. The corresponding sample cross-section made from this ring sample 
demonstrated that two cracks in prior P-phase were formed on the inside of the sample (that is, on that side 
which underwent a tensile stress). However, these cracks were not classified as through-wall cracks and, 
therefore, this test was not accompanied with the cladding failure. The next test was performed with a new 
ring sample. In this case, the test was terminated at the moment when the load drop not only stopped but the 
load began to increase again. The analysis of data obtained in this test (see Fig. 3.19) showed that the 
cladding failure was practically noted. Two cracks formed on the outer and inner sides of the cladding 
practically joined into the through-wall crack (though, in this cross-section a diminutive bridge of 
underformed material remained between two cracks). 

To supplement the data base necessary to work out the final conclusions on this issue, one more test with the 
E110 brittle ring sample was performed (see Fig. 3.20). 


3.22 












500 


400 


'300 - 


Load relief after the first load drop 


■3 

C 


' 200 - 


100 


2 3 4 

Relative displacement (%) 


400 


300 


-3200 


100 


Load relief after the second load drop 


3 4 

Relative displacement (%) 


# 25-8 













-1 







✓ 





shu 

O) 

t-i 

ft 

/own 

est 






KrOOOSe-lS-v 2 grf | 


#25-9 











Jr 









/ / 






/ 

/ 

^- 




— 

snui-aown 1 

iff test 

■ 

KrOOOie 1 4- v2 grf 

-1-1 


o 


□ 

-t-i 


vi 




| 




r -V ,v, 

Ofef-'VV' 


^ v * - 


v/Jl 


Ilu>;mi I 



^ jJfVS 

% j*a , v ^ . 

* £) i ^ 

-\‘rf Ziff \ 

ftf 

** - if i’ > 


/S 


7 g 



Fig. 3.20. The data base to characterize the mechanical behavior of the brittle ring sample 


The analysis of obtained data shows that the first load sharp drop is associated with the formation of the 
through-wall crack, that is, this point should be qualified as the cladding failure. The second (and the 
following) load drop indicates the formation of new through-wall cracks in the cladding up to the complete 
sample fragmentation. 

The generalization of all performed investigations allows to make the following conclusions: 

• the load displacement diagrams with the monotonic change in the load during the tests (samples with a 
high residual ductility) are processed using two approaches: 

- if the maximum relative displacement is higher than 60 % (see the upper left diagram in Fig. 3.14) it is 
considered that the sample failure takes place in the point of the sample compression termination (thus, 
the sample may remain actually unfailed); 


3.23 





















































- if the maximal relative displacement is less than 60 % (see the right diagram in Fig. 3.14) it is consid¬ 
ered that the sample failure takes place at the moment of the load sharp drop. 

• the load-displacement diagrams of brittle samples with the low ductility margin (see two lower diagrams 
in Fig. 3.14) are processed using the following method: 

- a local drop of the load (with the value not higher than several tens of newtons) is neglected; 

- the moment of the sample failure is determined in the point of the load sharp drop (with the value higher 
than 100-150 N). 


3.2.2.2.Ring tensile tests 


Simple ring specimens cut out of the oxidized El 10 cladding samples were used for ring tensile tests. Major 
provisions of these tests were described in [11, 19, 20]. It should be noted that scoping tests necessary to 
certify the procedure as applied to the oxidized cladding were not provided for in the present research pro¬ 
gram. Therefore, the obtained results should not be considered as a set of standard mechanical properties for 
the cladding material. These tests were aimed at the obtainment of a set of experimental data characterizing 
the tendency towards the cladding ductility reduction as a function of the ECR and at the determination of 
the zero ductility threshold (the critical ECR). The plastic component of relative displacement (S p i) was de¬ 
termined for each tested specimen in accordance with the procedure presented in Fig. 3.21. 

2500 

2000 

1500 

-a 

S3 

o 

_J 1000 


500 

0 

0 0.2 0.4 0.6 0.8 1 1.2 

Displacement (mm) 

Fig. 3.21. Processing of the load-displacement diagram obtained in the ring tensile test of a simple ring 

sample manufactured from the oxidized El 10 cladding tube 



3.2.2.3. Three-point bending tests 

Three-point bending tests of the oxidized El 10 claddings were performed in the universal test machine 
(1794Y-5) equipped with the device shown in Fig. 3.22. 

The oxidized El 10 cladding sample 80 mm long was placed on two cylindrical supports of the roller type 
5 mm in diameter. The distance between the supports was 70 mm. The third cylindrical roller bearing pro¬ 
vided for the impact of the bending load on the cladding sample with the rate of about 1 mm/min. All these 
tests were performed at 20 C. The load-displacement diagram was recorded during the tests. It should be 
noted that three-point bending tests of tubular samples were not of the standard type of mechanical tests. 
Therefore, the experience gained previously in similar tests was specially analyzed [21, 22]. The analysis 
allowed to formulate the following general requirements imposed on three-point bending tests: 

• a roller bearing should be employed as a support and a loading tool to minimize friction during the proc¬ 
ess of cladding bending; 

• the distance between supports must exceed the sample diameter not less than 8-10 times. 


3.24 

















Fig. 3.22. Three-point bending test apparatus 


The test apparatus illustrated in Fig. 3.22, complies completely with these requirements. Besides that, this 
approach is in a good agreement with the configuration of similar tests performed with the \15 allov samples 
[23]. 

The major provisions of the result processing procedure for three-point bending tests are shown in Fig. 3.23. 



Fig. 3.23. Schematic for the processing of the load-displacement diagram 

after three-point bending tests 


In accordance with the presented scheme, the residual deflection at failure (D p i) as a function of the ECR was 
used as a result of each three-point bending test. 


3.25 
















3.3. Discussion of test results and working out of preliminary conclusions 


3.3.1. Reference tests 

The goal of reference tests was to determine the dependence of cladding residual ductility on such 
parameters of the oxidation mode as heating and cooling rates, following which to develop recommendations 
concerning the oxidation conditions for other stages of the research program. The background of this stage of 
the work was the following: 

• the heating rate defines such processes as the phase transformations (a—»a+P—>P) and the diffusion 
redistribution of alloying and impurity elements in the cladding material, that in its turn predetermines 
the oxide type (monoclinic, tetragonal) formed on the cladding surface, his protective properties and, as a 
consequence, H 2 uptake by the cladding material; 

• the cooling rate determines specific features of such processes as transition of a high temperature P-phase 
into low temperature a’-phase as well as the size and distribution of solid hydrides in the prior P-phase 
of oxidized cladding. 

In spite of the fact that the above listed effects were subjected to the attention of many researchers 
performing similar works, the analysis showed that systematical studies to estimate these effects in the 
aggregate were not performed. Therefore, the appropriate reference tests were included in this research 
program in accordance with the following test approach: 

• cladding type: as-received El 10 tube; 

• oxidation type: double-sided; 

• isothermal oxidation: 1100 C; 

• temperature of mechanical tests: 20 C; 

• combination of heating and cooling rates during the oxidation: 

- F/F (fast heating at 25 C/s and fast cooling at 25 C/s); 

- F/Q (fast heating at 25 C/s and quench cooling at 200 C/s); 

- S/S (slow heating at 0.5 C/s and slow cooling at 0.5 C/s); 

- S/F (slow heating at 0.5 C/s and fast cooling at 25 C/s); 

- F/S (fast heating at 25 C/s and slow cooling at 0.5 C/s). 

The whole scope of test results obtained due to these tests is presented in Appendixes B, C, D; the organized 
test results are shown in Fig. 3.24. 


3.26 



Combination of heating and cooling rates: 


F/F, F/0 



F/S 



S/S 


S/F 




Fig. 3.24. Results of reference tests 


The comparative analysis of the mechanical behavior of oxidized El 10 claddings after reference tests al¬ 
lowed to reveal the following (see Fig. 3.25): 

• the higher the ductility margin in the oxidized cladding, the greater the extent for this margin to be a 
function of oxidation conditions; 

• the less the cladding heating rate and higher the cladding cooling rate, the less degree of oxidation for the 
proportional decrease of the cladding residual ductility; thus, the decrease of residual ductility down to 
40% (approximately 2 times in comparison with the unoxidized cladding) takes place in the following 
consecution: 

- S/F test mode: ECR=5%; 

- S/S test mode: ECR=6.3%; 

- F/F, F/Q test modes: ECR=7.1%; 

- F/S test mode: ECR=8.5%; 

• the mechanical behavior of oxidized claddings cooled at 25 C/s and 200 C/s (F and Q) does not differ; 

• the closer the cladding is approaching the brittle state, the less its mechanical behavior depends on the 
studied parameters of the oxidation scenario (see Fig. 3.26) because, apparently, the alloy type is that key 
factor on which the zero ductility threshold depends. 


3.27 








































H-- 1 ---‘-H-—--- 1 --- 1 

0 2 4 6 8 10 12 14 


ECR (%) as measured 

Fig. 3.25. The residual ductility of the El 10 cladding vs heating and cooling rates 





9.1 



8.2 


7.6 ! 

7.7 






S/F 

’ - /- 

F/F, F/Q 

F/S 


N® 


dZ 

U 

PQ 

-a 

<L> 

3 

cfl 

3 

O 


7 


Combination of oxidation modes 

Fig. 3.26. The zero ductility threshold of the El 10 cladding vs heating and cooling rates 


Thus, the performed reference tests have demonstrated that in spite of the wide range in the variation of heat¬ 
ing and cooling rates, the zero ductility thresholds of the oxidized El 10 cladding are in the narrow range of 
ECR (7.9-9.1%). Taking this circumstance into account as well as the fact that slow heating and cooling 
conditions (used in these tests) are not prototypical for the LOCA analysis and in addition the fact that the 
difference between the fast cooling and quench cooling was not observed, the following general recommen¬ 
dation was worked out on the basis of reference tests: to perform the oxidation of cladding samples at other 
stages of the research program at the F/F combination of heating and cooling rates. 


3.28 



















































3.3.2. The comparative analysis of El 10 and Zry-4 oxidation 
and mechanical behavior 


Two cladding samples of Zry-4 alloy were oxidized at 1100C and F/F, S/S combinations of heating and 
cooling rates. The oxidation mode with a slow heating and slow cooling was employed to reveal peculiar 
features in the oxidation of Zry-4 claddings under the most adverse test conditions. The general goals of tests 
with the Zry-4 cladding were as follows: 

• to verify the test procedures and test equipment using the previous test data obtained for the Zry-4 
claddings; 

• to clarify the major differences in the behavior of the El 10 and Zry-4 claddings. 

The initial characterization of the Zry-4 cladding and obtained test results are presented in Appendixes A-3. 
B. G. 


3.3.2.1. The verification of test procedures 


In spite of the fact that significant efforts have been put into the development and validation of test 
procedures used for this research, the practical experimental experience show's that the risk of systematical 
(unaccounted) errors is never too low. Therefore, the only reliable assurance of the fact that these errors are 
of the acceptable value may be found in the comparison of results obtained in different laboratories 
employing different experimental apparatus and experimental methods. 

The test programs performed in the following countries were selected to compare Zry-4 test results: 

• Russia, VNIINM [15]; 

• Germany, NC Rossendorf [24]; 

• Hungary, KFKI [25]; 

• Czech Republic, SCODA-UJP [26]; 

• France, Framatom [23]; 

• USA, ANL [27, 28]. 

The data published for the appropriate tests (table data and numbered graphical output) were used to develop 
the regression correlations for these six test sets (see Fig. 3.27). The obtained regression correlations are pre¬ 
sented in Fig. 3.28. 


3.29 


VNIINM 


— 

■ HhHBhHHHhH 



NC Rossendorf 



c 

<u 

£ 

<u 

CJ 

c3 

CL 

C/2 


OJ 

> 

~E 

<J 

Cci 


70 

60 

50 

40 

30 

20 

10 

0 


Op 

\ ° 







o 

c 

o 

900 C 

950-1000 C 

1050-1100 C 

— correlation 

- 





V 3 0 








Vc 





□ \ 

o 

L o 




□ 

0 

o 

o 


— 



o 

rd-bot-m-rry-l -v? 


0 


10 


15 


20 


25 



Measured ECR (%) 

mmm 




c 

0) 

E 

(U 

o 

— 

a, 

c/2 


cj 

> 

03 

CJ 

Q£i 


70 

60 

50 

40 

30 

20 

10 

0 


°\ 





8 \ 

% 








O 110 

OC 

elation 






- corr 













DO 





oS. 







rd framat- zry-l v ’ 


0 5 10 15 20 

Measured ECR (%) 


25 




KFKJ 


c 

OJ 

I 40 

c3 

Cl 

C/2 


<U 

> 

C3 


CJ 



5 10 15 20 

Measured ECR (%) 



60 


40 


5 50 

E 

CJ 
CJ 

_E2 

Oh 

~ 30 

•g 20 

e3 

o 

Ofii 10 


0 






-1 



O 1100 C 

- correlation 

- 


0 

\c 


p 







o \ 





8oV 

_ 




o ' 






\ 

\ 


rd-vrtil zn l-v 2 



O O* 

<b 



0 


i 10 15 

Measured ECR (%) 


20 


25 



Fig. 3.27. Summary of the ring compression test results performed in different laboratories 

with the Zry-4 oxidized cladding 


3.30 


























































































































































Fig. 3.28. The comparative test data characterizing the ductility of Zrv-4 oxidized cladding vs ECR 

The comparison of these data with the data obtained by RRC KI/RIAR allows to conclude that: 

• the previous VNIINM data [15] must be eliminated from the consideration and, probably, KFKI data 
[25], also (the motivation to this position will be discussed in the next sections of this Chapter); 

• the analysis of the data base confirms the conclusion developed in the section 3.3.1 of this report: “the 
higher the ductility margin in the oxidized cladding, the greater the extent for this margin to be a function 
of oxidation conditions”. Besides, in this case this margin is a function of the zircaloy type (because the 
test data base covers more than a twenty-year period, and, accordingly, it cannot be considered that 
characteristics of tested zircaloy types have been completely identical); 

• taking into account the preceding conclusion, different test data in the range 0-10% of the ECR were not 
compared; as for the presented range of the ECR (10-20%) then everybody can see that the RRC 
KI/RIAR test data are in the middle of the range for the experimental dispersion of results. 

One more proof of the RRC KI/RIAR test data representativity was performed basing of the comparison of 

data characterizing the Zry-4 oxidation kinetics (see Fig. 3.29). 


3.31 


























Fig. 3.29. The comparative data obtained in different laboratories 
to characterize the Zry-4 oxidation kinetics 


The analysis of comparative data shows that RRC KI/RIAR data are in a good agreement with the data ob¬ 
tained by other researchers. Consequently, the integral verification the RRC KI/RIAR test approach showed 
that possible systematical errors connecting with experimental and processing procedures do not exceed rea¬ 
sonable values. 


3.3.2.2. The comparison of the El 10 and Zry-4 oxidation and mechanical be¬ 
havior 


The comparison of the El 10 and Zry-4 residual ductility as a function of the ECR is presented in Fig. 3.30. 
70 


N° 

o^- 


O 

"O 

Is 

3 

T3 

<L> 

QC 


60 

50 

40 

30 

20 

10 

0 


8 


14 


16 












_ _ El 10 at F/F and F/Q combination 

of heating and cooling rates 



\ 





V 

\ 



+ 

a i i /i Lumumaiiui 

of heating and cooling ra 

tes 



\ 



— 

— 



\ 

\ 







\ 

\ 


+ # 




- 

— 

i 

\ 

-* - 




e-zr\4-en 


10 12 

Measured ECR (%) 

Fig. 3.30. The comparison of El 10 and Zry-4 cladding behavior in accordance 

with RRC KI/RIAR test data 


18 


* The characterization of the El 10 cladding and test results obtained at 1100 C and F/F, F/Q combinations of heating 
and cooling rates are presented in Appendixes A-3, B, D 


3.32 























































The analysis of these data allowed to state the following: 

• the zero ductility threshold of the El 10 cladding is 8.3% ECR; 

• a sharp decrease ot the El 10 cladding ductility (from 60% down to 0) occurs in a very narrow range of 
ECRs (6.7-8.3%) that indicates that some new physical phenomenon takes place in the mechanism of the 
El 10 alloy oxidation; 

• in accordance with the data presented in Fig. 3.28 the ductility decrease of the Zry-4 cladding is of a 
monotonic character, that is, one and the same mechanism accompanies the Zry-4 oxidation in the whole 
studied range of the ECRs; 

• the Zry-4 oxidized cladding has the residual ductility 14.3% (22.4% of the relative displacement) at 
11.5% of the ECR. 


To clarify the physical phenomena, which are responsible for the revealed differences in the El 10 and Zry-4 
oxidation behavior, the analysis of data characterizing the appearance of these cladding types was performed 
first of all (see Fig. 3.31). 


F/F 

Zry-4 

(#64) 


E110 

(#105) 








Fig. 3.31. Appearance of the El 10 and Zry-4 claddings after the oxidation at 11.3-11.8% ECR and 

F/F, S/S combinations of heating and cooling rates 


These first observations have shown the following: 

• the Zry-4 cladding is covered with the lustrous black ZrO: oxide in both variants of heating and cooling 
rates; 

• the El 10 cladding oxidized at F/F combination of heating and cooling rates has the spalled white oxide 
on the surface; 

• the E110 cladding oxidized at S/S combination of heating and cooling rates has the remains of the 
lustrous black oxide (outer layer) and very spalled oxide (inner layer). 

The noted peculiar features of the oxide state on the El 10 cladding surface indicated that the breakaway 
mechanism of the cladding oxidation took place. In accordance with the data presented in Fig. 3.32, the be¬ 
ginning of this type of oxidation is observed at 6.5% ECR in the form of local white spots on the cladding 
surface. On the ECR increase, the number of white spots is increased also right up to these white spots join¬ 
ing and producing the outer layer of the spalled oxide. 


3.33 























ECR=7.9% 


ECR=14% 


ECR=6.5% 


ECR=7.6% 


#46-9 


#65-10 


#82-7 


#110-4 


Fig. 3.32. The appearance in detail of the El 10 oxidized surface vs the ECR 


The additional illustration of this effect presented in Fig. 3.33 shows that two oxide layers are formed on the 
E110 cladding surface during the oxidation at the ECR higher than 7-8%. 



prior (3-Zr 


Internal layer 
of ZrO ; oxide 

a-Zr(O) layer 


External surface of oxidized sample 
-/ 


External layer 
ofZKX oxide 


Gap between 
oxide layers 


Fig. 3.33. Demonstration of two layers of Zr0 2 oxide on the outer surface of the El 10 
oxidized standard as-received tube using fractography results 


Thus, this stage of the analysis of the El 10 specific behavior allowed to determine that the breakaway oxida¬ 
tion occurs in the El 10 cladding where breakaway oxidation is not observed in the Zry-4 cladding at 1100 C 
[13]. Taking into account that the breakaway oxidation may be realized in the form of two different mecha¬ 
nisms (the uniform oxidation and nodular oxidation), additional examinations were performed to study these 
effects using the metallographic cross-section samples of oxidized claddings (see Appendixes C, D, G). 

In parallel with the RRC KI/RIAR studies, several cross-sections of El 10 oxidized claddings were prepared 
in the ANL. The interpretation of the ANL data was made by H.Chung (see Fig. 3.34). In accordance with 
his position, the nodular type of the breakaway oxidation was observed on the tested El 10 cladding samples. 
The combination of phenomena during the oxidation may be described in the following way [29]: 

• “initial development of white nodules; 

• coalescence and interconnection of nodules; 


3.34 



















• flake-of of gray-white continuous oxide layer”. 

This point of view was taken as the basis in our previous analysis of this program results [30]. But a more 
careful analysis ot this approach based on the metallographic studies performed by the RRC KI/RIAR has 
shown that there are some contradictions between this position and the experimental data. To demonstrate 
the revealed contradictions, the appearances of several oxidized El 10 cladding samples are presented in Fig. 
3.35. 



a The microeffects of 

ECR=6.5% , . ... 

breakaway oxidation 

#46 





white spots 

yf\ 

_ The macroeffects of 

UH2 

ECR=7.9% , . 

breakaway oxidation 

iipmi 




The heaving of large parts 

craters cracks 

ECR=10.5% => of oxide in the crater 


form, cracking of oxide 




The oxidation of previous 

, , n/ separated parts into the 

ECR=14% => /. ,. . , . 

stoichiometric white 

oxide 

white oxide black oxide 

UUO s' 





Fig. 3.35. The characterization of the oxidized El 10 claddings vs ECR 


Thus, if we are based on the concept of the nodular corrosion as the only mechanism determining the El 10 
cladding behavior as a function of the ECR then it should be expected that the cladding samples will be so 
much more white the higher is the ECR value. However, analyzing the appearance presented in Fig. 3.35 it 
can be seen that this thesis is confirmed while turning from the sample #46 to the sample #82 and is not con¬ 
firmed if to turn from the sample #82 to the sample #28 (as the major part of the sample #28 is covered by 
the lustrous black oxide having several separate big defects of the crater type, microcracks and the explicit 
tendency to the whole oxide layer exfoliation from the cladding surface). On the further ECR increase (the 
sample #110) it may be observed that these sections (separated earlier) of the tetragonal oxide layer located 
under the stoichiometric oxide are oxidized into the white stoichiometric oxide, while the oxide sections 
retaining the adhesion with the cladding material remain lustrous and black. 


3.35 






















To investigate these processes in more detail, metallographic samples of oxidized cladding were studied. The 
general tendency of processes, which take place in the El 10 cladding as a function of ECR, may be charac¬ 
terized using the data presented in Fig. 3.36. 

The consideration of these data confirms that: 


• the uniform oxide layer is formed on the El 10 cladding surface at the ECR<6.5%; 

• at the ECR=8.9% the formation of local parts is observed in which oxide is separated into individual 
layers due to cracking. In this case, at least the upper layers are of the stoichiometric oxide and are 
perceived as white spots on the visual inspection. 

Thus, the preliminary consideration of the nature of the El 10 high temperature oxidation allows to determine 
that the beginning of the El 10 oxidation process is accompanied by the initiation of the breakaway oxidation 
in small local points of the cladding surface, and this oxidation stage can be characterized as the nodular 
breakaway oxidation. But the analysis of a number of metallographic samples prepared to observe this proc¬ 
ess (see Appendixes C, D) has shown that on the ECR increase up to approximately 7% these primary effects 
remain on the oxide outer surface but are not developed into the depth. Therefore, a typical view of a devel¬ 
oped stage for the nodular breakaway oxidation presented for the comparison in Fig. 3.37 is not observed in 
the tested cladding samples. Moreover, a special set of metallographic samples prepared to reveal the space 
geometry of oxide layers (see Fig. 3.38) allows to reveal that quite an even boundary separates the a-Zr(O) 
from the outer and inner oxide layers. 




Polished 


ECR = 8.9 % 


H tort =l 110 ppm 


Average 

residual 

ductility: 

0.2 % 




** * . 
externa/ surface 


-V 


internal surface 


l 


Multu-layers 

ZrOj 

oxide 


Fig. 3.36. Visualization of Zr0 2 oxide behavior as a function of the ECR 
after the double-sided oxidation at 1100 C (F/F combination of heating and cooling rates) 

of E110 standard as-received tubes 


3.36 



























Fig. 3.37. The appearance and microstructure of Zr-Nb oxidized cladding after the operation with the 

surface boiling (RBMK cladding type) 



>/j] if M * 


2 ®! * i l ¥ ' *■**& 4 ' * ■- 


Suiflfjv? 

v *4 ( “ 


§ sm 

& I? 

" ^X'ih 

W» .;<*£> i >■■ 


V - r ■ -Jk * £*~. v- 

O.i'i v^S^, V 

<4 rsfe - t' fi , 

" ■ “ «§&, 


^ » .,i'. 


arf' 

/ 


/ 



'A V «*» 


/A ' 

*r> £ x*- .y 

^4 f •->* 

7./ v L«i *■ n 

k'/- ^ . ‘r.%> 



& \%£ J \ J ' Iff \ A W , ft 


Fig. 3.38. Angular variations of the El 10 standard as-received tube microstructure after a double¬ 
sided oxidation at 1100 C and 10% ECR (sample #17) 


3.37 



















The analysis of the El 10 breakaway oxidation nature will be continued in other sections of the report. But to 
complete the comparison of the El 10 and Zry-4 oxidation and mechanical behavior, several other important 
phenomena are considered in this paragraph: 

• oxidation rate; 

• the a-Zr(O) thickness; 

• oxygen absorption and distribution in the prior P-phase; 

• hydriding of prior P-phase. 

As for first position of the list, this research confirms that the El 10 oxidation rate is somewhat less then that 
of zircaloy. However, for the objectives of this research it may be considered that the El 10 and Zry-4 
oxidation kinetics are comparable. 


The next important factor influencing the mechanical behavior of the oxidized cladding is the a-Zr(O) phase 
thickness. A typical feature for the Zry-4 cladding oxidation behavior at high temperatures is the formation 
of three concentric layers in the oxidized cladding (see Fig. 3.39): Zr0 2 , a-Zr(O), prior-P phase. 



Fig. 3.39. The microstructure of the Zry-4 cladding (sample #64) oxidized at 11.5% ECR 

at F/F combination of heating and cooling rates 


In the Zr-Nb claddings (in contrast to Zry-4), the a-Zr(O) phase (in the form of needles) penetrate deeply 
into the prior-P phase producing an irregular boundary between these phases. Moreover, as a rule, the 
boundary line in these phases cannot be legibly determined (see Fig. 3.40). 


3.38 






















#42-5, Etched, ECR=8.9% 

Zr0 2 a-Zr(O) prior p 

i i 'i' 



Fig. 3.40. The microstructure of the El 10 cladding after a single-sided oxidation at 1100 C 


This peculiar feature of the Zr-Nb claddings is supplemented with the fact that the a-Zr(O) layer in the El 10 
cladding is noticeably thicker then that in the Zry-4 one (see Fig. 3.41). 



Fig. 3.41. The comparative data characterizing the Zr0 2 and a-Zr(O) thickness 


To clarify oxygen absorption by the El 10 cladding under oxidation conditions, several special investigations 
were performed including such methods as: 

• the measurement of microhardness across the cladding thickness; 

• the measurement of oxygen concentration across the cladding using the Auger spectrometry; 

• the determination of oxygen concentration across the cladding using the SEM (scanning electron 
microscopy) technique. 

The comparative data characterizing the microhardness and oxygen distributions in the cladding sample #41 
presented in Fig. 3.42 confirm the known thesis that the microhardness distribution corresponds to the oxy¬ 
gen distribution. Besides, both types of test data allow to support the following conclusion formulated earlier 
[24]: oxygen is uniformly distributed in the prior p-phase of the El 10 cladding in contrast to the Zry-4 clad¬ 
ding. 

To reveal the character of oxygen distribution in the Zr0 2 , a-Zr(O) and prior p-phase of the El 10 oxidized 
cladding, the SEM examinations of the reference cladding sample were performed using the 
XL 30 ESEM-TMP scanning electron microscope (FEI/Philips Electron Optics) equipped with INCA Energy 


3.39 



























300 (Oxford Instruments) EDX (energy dispersion x-ray analyzer). Those allowed to formulate several ob¬ 
servations concerning oxygen behavior in the El 10 claddings (see Fig. 3.43): 

• oxygen concentration in the a-Zr(O) phase of the niobium-bearing cladding is monotonically decreased 
from the metal-oxide interface down into the depth of a-Zr(O) layer; 

• non uniform distribution of oxygen concentration across the ZrC >2 layer has been revealed using the 
electron probe microanalysis, besides, the secondary electron image of zirconium dioxide has shown that 
two different types of ZrCE are observed on the surface of the reference cladding sample; 

• in accordance with the results of investigations performed to interpret similar processes taking place in 
zircaloy claddings during the base irradiation (see, for illustration [31]), the oxide of this type consists of 
the outer layer, in which the monoclinic phase (stoichiometric white oxide with the maximum oxygen 
concentration) of zirconium dioxide prevails, and of the inner layer which represents the tetragonal 
modification (understoichiometric oxide with lower oxygen concentration). 


3.40 


white spots 


Appearance of the 
E110 cladding 
after the oxidation 
at 8.2% ECR 


A-A: The cross-section of 
metallographic sample 



Microstructure of 
the oxidized 
cladding (etched) 


Microhardness 
across the cladding 
thickness 


=Ci 

xr. 

\r. 

O 


1000 


800 


600 


'= 400 


200 


#41 







• 

• • 

• • 

• 






• 

• 

• 

• 

• 

• 

• 

• 

• 

* • 






* 

• • 

• 

• 

M 

• 

.. .* 

r. 

r •§ • 

• 

•» ■>./ 

• 

9 • 

• • 

# nr 

• • • 

.* 

• 




* . 




7 7 7 7 7 

7 7 7 7 7 7 

7 7 7 7 7~ 

7 JK4 / / 

7777777777777777 

' unoxidized cladding 


0.1 


0.2 


0.3 


0.4 


0.5 


0.6 


0.7 


a-Zr(0) prior p a-ZrtO) thickness (mm) 


The oxygen distri¬ 
bution across the 
cladding deter¬ 
mined by the Au¬ 
ger analysis 



a-ZrtO) prior P a-Zr(0) thickness (mm) 


Fig. 3.42. The appearance of microstructure and characterization of oxygen distribution 
in the El 10 cladding after the double-sided oxidation at 1100 C 


3.41 
































Secondary electron image of the 
sample #41-4 with four tracks of 
oxygen concentration 
measurements 


EPMA results (Electron probe x- 
ray analysis) characterizing the 
oxygen distribution 

Comments: 

1. Averaged data on four tracks is 
presented 

2. Measurements at the low oxygen 
concentration in the prior P-phase are more 
qualitative than quantitative. The 
systematic error in the measured data was 
caused by the parasitic oxidation of the 
sample surface 


II1P& 



V® 


c 

_o 

13 

■ 4 —* 

a 

<o 

o 

c 

o 

O 


40.0 


30.0 


20.0 


10.0 


0.0 


ZrO z 

4 

a-ZrO 

i 

ex-pZr 

4 

"V 

















& -Q 




20 


40 


60 


80 


100 


Distance (pm) 


Oxygen EDX dot map 



Fig. 3.43. The oxygen distribution in the El 10 cladding (1100 C, 8.2% ECR) 
in accordance with results of SEM examinations 


Obtained experimental data prompt to the formulation of the conclusion that general differences in the Zry-4 
and the El 10 ductility margin as a function of the ECR may be explained by the differences in the oxygen 
absorption and distribution. However, a more careful analysis shows that this explanation developed on the 
basis of previous described data contradicts the results presented in Fig. 3.44. 

So, two E110 claddings one of which is brittle and another one is ductile have practically the same 
microhardness (oxygen) distributions across the cladding material. In this connection the question arises: is 
the oxygen concentration in prior P-phase the only criterion determining the ductility margin of the oxidized 
cladding? 


3.42 








































The oxidized sample at the zero ductility threshold 
_(residual ductility- is 0.0%) 


The oxidized sample with a high margin of residual 
ductilitv (residual ductilitv is 57.7%) 


ECR-8.2% 


1000 


800 




■es 6oo 


o 

-5 


3 400 


200 


#41 






• 

>• 

• • 

m 

• 





• 

• 

• 

• 

•« 





•• 

• 

m 

• 

7 •* * 

T- 


» 

- -w 





V • 




7 7 7 

/////////// 

unoxidized cladding 

/ / / 

/ / / / 


0.1 0.2 0.3 0.4 0.5 0.6 

q-Zr(Q) prior ft q-Zr(Q) thickness (mm) 


0.7 


ECR=7.0% 


1000 


800 


600 


400- 


200 - 


• 

• 






#47 

m 






• 

• 






• 

• 

• 

• 

• 


\ 

• 

•. •• 

. • *.« 

: •• 

• 

M 








7 7 7 2 

*777 

77777777777 

unoxidized cladding 

7 7 7 

7 7 7' 7 


0.1 0.2 0.3 0.4 0.5 0.6 

q-Zr(O) / prior {3 / a-Zr(O) thickness 


0.7 


Fig. 3.44. The comparison of microhardness distributions for the brittle 

and ductile El 10 oxidized cladding 


The previous classical investigations performed with the Zry-4 claddings have shown that if the breakaway 
oxidation takes place then two competing processes determine the reduction of ductility margin of the 
oxidized cladding [13, 32, 33]: 

• oxygen absorption by the prior [3-phase; 

• hydrogen uptake by the prior (3-phase. 

The significance of the hydrogen effect was later confirmed by the studies performed with the El 10 cladding 
[17, 18, 24, 25, 26, 27], In accordance with current conception, the effect manifests itself in those cases when 
the hydrogen concentration in the prior (3-phase becomes higher than the hydrogen solubility limit in the 
zirconium. Taking into account that this limit is a strong function of the temperature, the following physical 
phenomena occur during the cooling phase with the oxidized cladding absorbing a significant hydrogen 
portion at the oxidation: 

• the formation of solid hydrides at low temperatures; 

• the decrease of residual ductility margin caused by hydriding of the prior P-phase. 

Obviously, the process of solid hydrides formation is the function of cooling rates (the size of hydrides and, 
possibly, the orientation of hydrides). These physical phenomena allow to understand the sensitivity of the 
E110 residual ductility to the cooling rates considered earlier, in spite of the specific response of the El 10 
cladding on the slow cooling (the improvement of mechanical behavior) which requires the additional 
analysis. 

The quantitative analysis of hydrogen absorption by the El 10 claddings will be presented in the next section 
of the report. But the qualitative consideration of the physical phenomena corresponding to this process is 
performed in the completion of comparative studies of Zry-4 and El 10 claddings. 

As it was noted in the comments to Fig. 3.42, the brittle El 10 cladding has two different layers of zirconium 
dioxide: 

• white porous monoclinic oxide; 

• black protective tetragonal oxide. 

The characterization of these two layers of oxide may be also added with the data obtained due to the fracto- 
graphy observ ations (the SEM studies of fracture surfaces in the reference cladding sample) presented in Fig. 
3.45. 


3.43 































Fig. 3.45. Demonstration of the morphology of Zr0 2 layers in the El 10 oxidized cladding 

(sample #41-4, 1100 C, 8.2% ECR) 

To understand a general succession of phenomena that may be responsible for the El 10 specific behavior, 
the following should be preliminary pointed out: 

• stabilization of the tetragonal protective oxide layer on the cladding surface is explained, as a rule, by 
compressive stresses available which are the consequence of the fact that the oxide volume is 
significantly larger than the initial zirconium volume (in accordance with Pilling-Bedworth ratio) and 
then the modification of the surface energy balance in the separate Zr0 2 grains; 

• besides, the stabilization of the tetragonal phase (slowing-down of the transition of this phase into 
monoclinic phase) or the improvement of its protective properties is determined also by the behavior of 
the alloying elements and secondary precipitates in the Zr matrix. 

Thus, it is known that the improvement of the corrosion resistance of Zr-Sn alloys was obtained due to the 
addition of such transition metals as Fe, Cr, Ni at very low concentrations (<0.5%) [31]. A general 
explanation of this effect is the following: the presence of liquid unoxidized Sn and unoxidized Fe, Cr, Ni 
precipitates in the zirconium dioxide prevents the hydrogen diffusion and stabilizes the Zr0 2 tetragonal 
phase. It is also important to point out that tin oxides exist in the tetragonal form only that also leads to the 
stabilization of tetragonal Zr0 2 . As for the niobium-bearing alloys, general considerations on this issue may 
be formulated in the following way: 

• in accordance with results of the diffusion of niobium in the cladding surface layers and its oxidation into 
Nb 2 0 5 may result in destabilization of Zr0 2 layer [29]; 

• the reason for this destabilization may be associated with the following circumstances: 

- if niobium is distributed in Zr0 2 non-uniformly (that is, zirconium dioxide represents a heterogeneous 
structure containing niobium enriched areas) then, during oxidation of this heterogeneous structure, high 
volume stresses will occur due to the difference in densities of Zr0 2 and Nb 2 0 5 , as well as due to 
stresses caused by the phase transition; 

- as it is known, the phase transition takes place practically instantly, therefore, it may be assumed that at 
that moment when the thickness of the oxide film achieves a definite value the phase transition occurs 
and the oxide film goes away from the oxide metal interface; 

- thus, it becomes evident that all precipitates in the Zr0 2 that are oxidized slowly (such as Sn, Fe) will 
retain the understoichiometric state of oxide and will correspondingly hinder the sharp phase transition 
by the stabilization of the tetragonal phase and vice versa. 

To clarify the niobium behavior in the El 10 alloy at the oxidation, several types of SEM examinations were 
performed. The results of these examinations are presented in Fig. 3.46. 


3.44 











Secondary electron image of the 
sample #41 -4 with four tracks of 
niobium concentration 
measurements 


Niobium WDX (wave length 
dispersive x-ray analysis) dot 
map 




3.45 





































































Fig. 3.46. The niobium distribution in the El 10 cladding (1100 C, 8.2% ECR) 
in accordance with results of SEM examinations 

The analysis of obtained data has shown that: 

• in accordance with the niobium WDX dot map, prolonged niobium enriched areas (white areas) alternate 
with niobium depleted areas (dark areas) in the a-Zr(O) phase of the El 10 cladding; 

• the oxidation of a-Zr(O) into Zr0 2 results in the fact that these niobium enriched areas decay into 
separate niobium enriched points in which niobium concentration is still higher; thus, studied oxide 
represents heterogeneous Zr0 2 with disseminations of niobium enriched local areas; 

• niobium distributions into four tracks obtained due to the EPMA confirm these observations; the 
irregular niobium distribution in Zr0 2 and a-Zr(O) reflects the niobium concentrations in the niobium 
enriched areas (up to 2.3%) and in the niobium depleted areas (0%); 

• the dot map image and EPMA measurements show that the prior (3-phase is characterized by quite a 
uniform distributions of niobium precipitates (white point on the dot map) in the a’-Zr matrix with the 
basic concentration in the alloy (1%). 

Thus, the experimental data allow to understand the nature of the irregular boundary between a-Zr(O) and 
prior (3-phases in the zirconium niobium alloys. The segregation of niobium with the formation of sequenced 
in the radial direction areas with different Nb concentration leads to the fact that niobium enriched areas 
transform into the a-Zr(O) phase at a higher oxygen concentration than the neighboring areas with lower 
niobium concentration. This effect determines the irregular boundary front between a-Zr(O) and prior 
P-phases in these alloys. 

The performed cycle of studies allowed to return to the consideration of the postponed earlier issue concern¬ 
ing hydrogen absorption by the prior P-phase of the El 10 cladding. One of the potential causes that may lead 
to hydrogen penetration through the oxide-metal boundary has already been named. This cause is the crack¬ 
ing and spallation of oxide due to the tetragonal-monoclinic phase transition in the heterogeneous oxide. 

The following specific features of the niobium-bearing cladding discussed above determine other possible 
causes of hydrogen penetration: 

• viens of Nb-rich P-phase in the a-Zr(O) phase represent some sort of channels providing hydrogen 
diffusion in the prior P-phase matrix; 

• hydrogen concentration diffusion is also stimulated by the fact that hydrogen solubility is negligibly low 
in the a-Zr(O) phase and the surface through which hydrogen diffuses is very large (due to the fact that 
a-Zr(O) and P-phases have s complicated irregular boundary); 

• the preferred radial direction of the a-Zr(O) grain revealed by metallographic examinations also 
facilitates hydrogen diffusion along grain boundaries into the cladding depth. 

Numerical investigations, which were performed earlier with different zirconium alloys, allowed to establish 
the following principles of hydrogen embrittlement: 

• during a high temperature oxidation, hydrogen absorbed by the p-phase is in the solid solution as H 2 
solubility in the P-phase is very high under these temperatures; 

• during the cooling phase, the precipitation of solid hydrides occurs because hydrogen solubility is a 
strong function of the temperature, a sharp decrease in hydrogen solubility takes place at the temperature 
less than 550 C; the solubility limit becomes very low at temperatures less than 100-150 C; 

• investigations performed with Zry-4 claddings have shown that the number of hydrides in the prior 
P-phase is increased up to the hydrogen concentration of 300^400 ppm, at a higher level of hydrogen 
concentration, the number of hydrides is not increased but the size of these precipitates is enlarged; this 
effect leads to the initiation of internal stresses and deformation of the cladding matrix; 

• the oxidized cladding embrittlement is not a function of hydrogen content only, the embrittlement is a 
function of hydrogen content, hydrides morphology and hydrides orientation; 


3.46 


• the most negative effect takes place in that case when hydrides form long chains along which a crack 
may extend. 

To lind out to what extent the hydrogen embrittlement problem is important for El 10 samples tested in this 
program, two types of investigations were performed: 

1. The SEM examinations provided the visualization of hydrides. 

2. The hydrogen content was measured in the oxidized cladding samples. 

The results of the first of these investigations are presented in Fig. 3.47. 


BSE image (back 
_ scattered elec¬ 
trons) of the oxi¬ 
dized cladding 
(#41-1, 1100 C, 
ECR=8.2%) 



Fig. 3.47. The SEM micrograph of the E110 oxidized cladding 


The backscattered electron (BSE) image of the oxidized cladding may be generally interpreted taking into 
account the following peculiar features of this method: 

• the less is the atomic number of a chemical element in the scanning point the darker is the point on the 
BSE image; 

• thus, the areas with high density look as the white ones at the micrograph (white areas), the areas with 
low density (such as hydrides) look as the black ones. 

In accordance with this pattern of the BSE image interpretation, it may be assumed that dark lines scattered 
about the white field of the Zr prior P-phase represent hydrides (though one cannot also exclude that a part of 
these lines may indicate the a-Zr(O) phase located along the boundaries of P-Zr grains as this phase density 
is lower than that of Zr). 

To extend the data base necessary for the analysis, the micrographs of one and the same sample obtained due 
to the optical microscopy are presented in Fig. 3.48: 

• the etched structure if the sample allow to see the distribution of a-Zr(0) phase inside the prior P-phase; 

• the polished structure of the sample shows the set of solid precipitates in Zr matrix, which could be in¬ 
terpreted as hydides. 

Comparison of the BSE image and optical image of polished cladding allow to assume that the prior p-phase 
of this cladding contains the plate type hydrides. Much of these hydrides are oriented in radial direction and 
therefore are more critical from viewpoint of the crack propagation during the mechanical loading. 

The measurement of the hydrogen contents in the El 10 and Zry-4 cladding samples shows that: 

• hydrogen content in the El 10 sample oxidized at 8.2% ECR was 1130 ppm; 

• the hydrogen content in the Zry-4 samples tested at 11.3-11.5% ECR and S/S and F/F combinations of 
heating and cooling rates was 34-37 ppm. 


3.47 













These measurements confirmed that the embrittlement of Zry-4 cladding is caused by the oxygen absorption 
in the prior P-phase. The embrittlement of the El 10 cladding is provided by the oxygen absorption and hy¬ 
drogen absorption in the prior P-phase. 








SA ,6 


ii 

' 

A 


I V 

OK 


#41-5 ECR=8.2% 


Etched 


#41-5 ECR=8.2% 


Polished 


Fig. 3.48. The optical microstructure of the El 10 oxidized sample with hydrides in the prior p-phase 


3.3.3. Determination of sensitivity of the El 10 cladding embrittlement to the 
oxidation type and the characterization of comparative behavior 
of El 10 and M5 claddings 

The previous stages of this research program revealed that the El 10 cladding has the inclination to the early 
breakaway oxidation, and to the increased hydrogen pickup, which facts lead to the zero ductility threshold 
of the E110 cladding of 8.3% ECR at 1100 C with double-sided oxidation. In this connection, the following 
natural question was formulated after the analysis of obtained data: do these results characterize the whole 
family of niobium-bearing claddings or does this effect take place in the El 10 cladding only? It is evident, 
that the comparison of the El 10 cladding behavior with the behavior of any other cladding with a close 
chemical composition might become the best method to find the answer to the formulated question. The 
appropriate analysis has shown that the M5 cladding [34] is an ideal partner to perform this comparison as 
both claddings (El 10 and M5) are fabricated from the alloys practically similar in the chemical composition: 
Zr-l%Nb (though it should be noted that they differ in oxygen concentration in the alloy). 

However, the consideration of published data characterizing the mechanical behavior of the M5 oxidized 
cladding showed that the appropriate studies were performed employing a single-sided oxidation [23, 35, 
36], This circumstance became the major incentive to perform sensitivity studies and to develop the com¬ 
parative data on the El 10 cladding behavior as a function of the oxidation type. The results of El 10 single¬ 
sided tests are presented Appendixes B and E. 

The obtained data have shown that (see Fig. 3.49): the pronounced indications of the breakaway oxidation 
appear in the case of a single-sided oxidation at lower ECR than that on the double-sided oxidation. This fact 
confirms again the formulated earlier thesis that the tetragonal oxide transition to the monoclinic oxide takes 
place on some critical oxide thickness. Taking into account that the oxide thickness at the single-sided 
oxidation is higher than that at double-sided oxidation (with the same ECR), the breakaway oxidation occurs 
at lower ECR. 


3.48 





















Double-sided 
oxidation => 

(ECR=6.5%) 

#46 


#46-5 


Etched 


#54-5 


Etched 


Fig. 3.49. The comparison of the El 10 appearance and microstructure 
after single-sided and double-sided oxidation at 1100 C 


The comparison of the El 10 oxidized cladding mechanical behavior presented in Fig. 3.50 allows to con¬ 
clude that the El 10 cladding demonstrates the tendency to the increase of the residual ductility margin at the 
single-sided oxidation (in comparison with the double-sided oxidation) but this difference is not so great. 



Measured ECR (%) 


Fig. 3.50. Comparative data characterizing the El 10 residual ductility 

as a function of the oxidation ty pe 


The E110 data base obtained due to the tests under the single-sided oxidation conditions allowed to perform 
a direct comparison of the mechanical behavior of the El 10 and M5 claddings on the basis of ring compres¬ 
sion tests and three-point bending tests at 20 C. To develop the comparative data, the M5 test results pub¬ 
lished in Reference 23 were used. Three-point bending test results are compared in this section and the sec¬ 
tion 3.3.7 of the report. The first set of comparative data organized in Fig. 3.51 allows to formulate several 
interesting observations listed in Fig. 3.51. 


3.49 



















































Ring compression tests 
with the El 10 claddin 


General observations: 

1. All tested El 10 claddings had 
approximately the same effective 
elastic modulus 

2. The fracture load at the zero duc¬ 
tility threshold (18.5 mg/cm") is 
corresponds to the “yield” point of 
load-displacement diagrams of par¬ 
tially ductile E110 samples (9.8- 
12.3 mg/cm : ) 


General observations: 

1. The claddings oxidized at 8.1- 
13.6 mg/cnr) had the same effective 
elastic modulus 

2. The elastic modulus of the sample 
oxidized at 22.9 mg/cm is lower 
than that of other samples 

3. The maximum load of 
22.9 mg/cm 2 sample is lower signifi- 
cantlv than the “yield” point 


General observations: 

1. All tested samples had approxi¬ 
mately the same elastic modulus 

2. Taking into account that the frac¬ 
ture load of 22.9 mg/cm 2 sample is 
significantly lower than the “yield” 
points of other samples the zero duc¬ 
tility threshold of M5 cladding is 
located in the range of 13.6- 
22.9 mg/cm 2 




Displacement (mm) 




Displacement (mm) 


~o 

03 

O 

►J 


2500 

2000 

1500 

1000 

500 

0 


M5 1231, nuni 

beret 

1 by RRC 

KI 


/ 

/ 


8.1 mg 

r /cm 2 

mg/cm 2 






*****»,* 




13.6 mg/cm 2 

_ 




22.9 r 

ng/cm 

i 2 

— 


— 

J-m5-bend-v2 


0 


2 4 6 

Displacement (mm) 


Fig. 3.51. The data base characterizing the mechanical behavior of the El 10 and M5 claddings 

after a single-sided oxidation at 1100 C 


The second set of El 10 and M5 comparative data presented in Fig. 3.52 allows to make the following con¬ 
clusions: 

• • 1 

• at the low oxidation (8.1-9.8 mg/cnr) the El 10 cladding demonstrates a higher level of residual ductility 
and lower strength properties than the M5 cladding. Apparently, this may be explained by the effect of 
different initial oxygen concentration in this two claddings; 


3.50 



























































• at the middle oxidation (12.3-13.6 mg/cm') both claddings demonstrate a very similar behavior at the 
initial phase ot the mechanical loading (the same elastic modulus and the similar level of the maximum 
load (800 and 720 N for the M5 and El 10 claddings respectively) but the M5 cladding has a noticeably 
higher margin of residual ductility; 

• at the high oxidation (18.5-22.9 mg/cm") the El 10 cladding demonstrates the transition from partially 
ductile state (14.8 mg/cm') to fully brittle state (18.5 mg/cm"); in this case, taking into account that the 
maximum load tor this sample corresponds to the “yield” point, it may be considered that this sample 
was tested at the zero ductility threshold; as for the M5 cladding, the load—displacement diagram of the 
appropriate sample (22.9 mg/cnr) shows that on the one hand, this sample has the margin of residual 
ductility (because the yield area is observed) but on the other hand, this sample had the maximum load 
that was significantly lower than the yield point, the effective elastic modulus of this sample was signifi¬ 
cantly lower than that ot other samples, and, besides three-point bending test showed that this sample 
was fully brittle; that is why our opinion is that the zero ductility threshold of the M5 cladding is lower 
than 22.9 mg/cm'. 



Fig. 3.52. The comparison of the El 10 and M5 cladding mechanical behavior after the single-sided 
oxidation at 1100 C in accordance with the ring compression test results at 20 C 


3.51 






































Nevertheless, the comparison of the El 10 and M5 cladding mechanical behavior has shown that the M5 
cladding has a higher margin of residual ductility at the oxidation level in the weight gain range of 13- 
20 mg/cm . Besides, the following should be pointed out: 

• the El 10 cladding is embrittled in accordance with the mechanism of the breakaway oxidation; 

• the M5 cladding oxidation takes place without any indication of the breakaway oxidation (in accordance 
with the test data presented in [23], the hydrogen concentration in the oxidized samples was very low and 
the oxide was lustrous and black). 

Thus, the results of this stage of research have shown that there is some cause that determines a specific 
behavior of the El 10 cladding. To determine more accurately the temperature range of this El 10 specific 
behavior, a special subprogram for the oxidation and mechanical tests was developed. The results of this 
subprogram are analyzed in the next section of the report. 


3.3.4. The evaluation of the El 10 oxidation and mechanical behavior 
as a function of oxidation temperature 


The results of investigations performed with the El 10 cladding on the oxidation at 1100 C have shown that 
this cladding has the tendency to the breakaway oxidation and the embrittlement caused by the oxygen and 
hydrogen absorption in the high temperature P-phase. Moreover, tests with the El 10 cladding performed at 
slow heating have demonstrated that the breakaway oxidation effects become still more pronounced. This 
may be explained by the fact that the initial phase of oxidation took place under the temperatures that were 
significantly lower than 1100C. These facts are also confirmed by the results of studies performed by 
V.Vrtilkova. She has shown that if the El 10 cladding is heated up to 1100C in the inert gas (argon) 
atmosphere and then is oxidized in water steam, the effect of early breakaway oxidation vanishes [10]. 

The analysis of these results leads to the conclusion that the most unfavorable test mode for the El 10 alloy is 
the oxidation in the area of the alloy a+P-phase. As for the temperature range at which the El 10 cladding has 
the a-phase, then numerous tests as well as the operation experience demonstrate that the El 10 claddings are 
not susceptible to the breakaway oxidation that is confirmed by the results presented in Fig. 3.53 [37, 38]. 


Typical appearance of the 
E110 commercial clad¬ 
ding after irradiation at 
60 MW d/kg U and rep¬ 
resentative parameters of 
the appropriate claddings: 

• oxide thickness 
10 pm 

• hydrogen content 
50 ppm 




■i" ^ 


S 


> - . V .•••* . • ; 



Mechanical properties of 
the El 10 commercial 
irradiated cladding [19] 


10 


8 - 


c 

o 


OO 6 

c 

_o 

<U A 


C 

D 


-- 

8 

o 



Ring specimens (transverse) 
unirradiated 

irradiated 

-- 

• 

o " -* 

• 


o 

• 

• 

• 

_ _ • 

• 

• • 

—w 

4^ • 

— 

. 


• 




- 

8 m 




# • 





— 



— 


ue 04nn wo ax-v2.grf\ 


200 


400 


600 800 
_ Temperature (K) 


1000 


1200 


1400 


Fig. 3.53. The characterization of the El 10 commercial cladding after irradiation and oxidation 

in the a-phase 


3.52 










































To clarify the details for the oxidation and mechanical behavior of the El 10 cladding as a function of 
temperature, special tests were performed in the range of 800-1000 C. Besides, to replenish the data base 
with the results characterizing the El 10 behavior at the maximum temperature typical of the design basis 
accident area, several tests were performed at 1200 C. The results of these tests are presented in Appendixes 
B and D of the report. 

But before analyzing the results of appropriate tests, it is useful to consider peculiar features of the a to 
P-phase transformation as applied to Zr-l%Nb alloys. The comparative data characterizing this phenomenon 
in the El 10 and M5 alloys are presented in Fig. 3.54. 



o 

rz 


o 

s. 

r3 

C_ 

C3 

O 



Temperature (C) 

Fig. 3.54. The characterization of the allotropic phase transformation in the El 10 and M5 alloys 


The analysis of presented data on the M5 alloy shows that the increase of heating rate leads to the 
displacement of the zirconium a-P phase area towards higher temperatures; with heating rates typical of the 
fast mode (F) of these tests this effect will be still noticeable. The studies performed in the RIAR during the 
recent years to understand the behavior of VVER fuel rods with the El 10 claddings under RIA conditions 
showed that in addition to the effects of heating rates as applied to the zirconium matrix the behavior of 
niobium should be also taken into account. So, the appropriate analysis shows that at the fast heating, the 
diffusion processes associated with the dissolution of P-Nb precipitates may take place not in full and, 
consequently, dissolved niobium atoms will localize close to those areas in which P-Nb precipitates are 
situated. In other words, the areas of niobium oversaturated solid solution may be generated in the zirconium 
matrix under these conditions. It has been already said before that such niobium-enriched areas may be 
responsible for the increased hydrogen absorption and other effects being of importance for understanding of 
the El 10 alloy peculiar features. 

The analysis of the El 10 alloy real behavior in the temperature range 800-1100 C may be started from the 
consideration of data characterizing the appearance and microstructure of oxidized samples (see Fig. 3.55, 
Fig. 3.56). The results of this analysis may be formulated in the following way: 

• the earliest appearance of the pronounced indications for the breakaway oxidation at the low level of 
ECRs is noted in the temperature range 900-1000 C; 

• at the middle level of ECRs the breakaway oxidation is visually observed under all studied temperatures 
(800-1100 C), however, the strongest indication of appropriate effects was noted at 900-1000 C; 

• at the high level of ECR the most striking demonstration of the breakaway oxidation was noted at the 
temperature of 950 C, in which connection the microstructure analysis showed that the effect of the 
oxide foliation was observed on samples tested at 800-950 C, but this effect manifested itself most 
obviously under the temperatures 900-950 C. 


3.53 























800 C 


900 C 


1000 C 


1100 c 


ECR=3.4% 


ECR=3.9% 


ECR-5.7% 


ECR=6.5% 



#119 



#46 



800 C 


900 C 


1000 C 


1100 c 


ECR=8.6% 


ECR=6.7% 



#144 



#131 


ECR=7.6% 

#45 



ECR=7.9% 

#82 







1100C ECR-10.5% 

#28 


800 C 


ECR=11% 


#132 


900 C 


=> ECR=12.3% 

#142 


1000 c 


Fig. 3.55. Appearances of the El 10 cladding after the double-sided oxidation at 800-1100 C and F/F 

combination of heating and cooling rates 


3.54 









































800 C 


Etched 


Etched 


#132- 

ECR-1 



900 C 






Etched 


Etched 


Etched 


Etched 


#141-5 

ECR=11.2% 


Fig. 3.56. The microstructure of the El 10 cladding after a double-sided oxidation at 800-950 C and 

F/F combination of heating and cooling rates 


3.55 


















































To understand consequences associated with the availability of stated above effects of the El 10 oxidation, 
the results of mechanical tests were specially organized (see Fig. 3.57). 


70 1 

An . 


+ 4-A 








OU 

Co 

so - 



*** ^ «■» 

t ° 

\ 









+ 800 C 

a 900 C 

O 950 C 


w jU 

1 40- 

3 

T3 




1 

t 

\ 

\ 








'A 

» 

A 




o 100 

oc 

DC 


G3 

3 

T3 

cZ in - 




A 

t 

i 

i 




1 1U 







<u 

& 

i n - 



o 

t 

o \ 

° ' 

+ 

+ 

+ 





1 u 

0- 




o 

—-e-n 

\ 

\ 

\ 

H - el 

+ 

+ 

b - 0 < 

e- ±4- 

r~0- 

e-all-950-en 


0 2 4 6 8 10 12 14 16 18 


Measured ECR (%) 

Fig. 3.57. The data base characterizing the residual ductility of the El 10 cladding 
as a function of the ECR and oxidation temperature 

In accordance with the data presented in Fig. 3.57, the following additional comments may be made: 

• the zero ductility threshold of the El 10 claddings oxidized at 800 C is noticeably higher than that in the 
reference oxidation mode (1100 C); 

• the ECR of the El 10 claddings oxidized at 900 C is approximately the same as the 1100-degree 
threshold, but the time threshold is much higher at 900 C than that at 1100 C; it is outside a practical 
LOCA; 

• the zero ductility threshold at 1000 C is possibly a little lower than the 1100-degree threshold. 

These observations are in some contradictions with the observations formulated according to the analysis of 
the cladding appearances and microstructures; but before we go on to the consideration of possible causes for 
these contradictions, it is useful to add the results of hydrogen concentration measurements in the oxidized 
claddings to the data base. 

The systematic analysis of the data presented in Fig. 3.58 allows to reveal the following regularities in the 
E110 cladding behavior as a function of the ECR at 1100 C: 

• the mechanical behavior of oxidized cladding is characterized with the very high margin of residual duc¬ 
tility in the ECR range approximately up to 7%; 

• the cladding hydrogen content is very low in the same range of ECRs; 

• the sharp decrease of residual ductility in the ECR range 7-8.3% corresponded with the sharp increase of 
hydrogen content in the El 10 cladding; 

• the zero ductility threshold of the El 10 cladding is corresponded with the hydrogen concentration of 
400 ppm approximately. 

Thus, the data base characterizing the hydrogen content in the El 10 cladding after the oxidation at 1100 C 
confirms formulated earlier assumptions concerning the fact that the embrittlement of the El 10 cladding is a 
sum of the oxygen and hydrogen embrittlement. Besides, it should be noted that this critical value of the 
hydrogen content in the cladding (400 ppm) is in a good correlation with the above mentioned critical value 
of the hydrogen content, after which the increase of the hydrides volume and internal stresses in the cladding 
were observed. 

Specific features for hydrogen absorption by the El 10 cladding as a function of the oxidation temperature 
may be discussed using the data given in Fig. 3.59. 


3.56 







































~z 

■s. 



Measured ECR (%) 


1200 


9 800 4 

EL 
c. 


C-) 

1 400 


0 


E110 

Double-sided 
T ox =l100C 

FT, F/Q 


o 

o 





i 

i 

i 

Q 

O l 
l 

1 O 

1 

1 

1 

1 

1 


G 

i 

o° 

» 

l 


0 4 8 

Measured ECR (%) 


12 


Fig. 3.58. The El 10 residual ductility and hydrogen concentration as a function of the ECR after a 
double-sided oxidation at 1100 C and F/F, F/Q combinations of heating and cooling rates 


c_ 


CJ 


tJ 
CXj 



4000 


3000 


2000 


1000 


0 







A 


+ 800 C 

A 900 C 

o 950 C 

-1100C 





o 



/ 






±A, 

A 

/ 

/ + 

' + 

h2-en-all-v2 

-1 


0 


4 


8 


12 


16 


Measured ECR (%) 

Fig. 3.59. The hydrogen content in the El 10 cladding as a function of the ECR and temperature 

after a double-sided oxidation 


A part of obtained data is in a good agreement with the results of mechanical tests presented in Fig. 3.57, so: 

• the hydrogen absorption decrease at 800 C leads to the improvement of the El 10 mechanical behavior 
and to the increase of the cladding zero ductility threshold up to the ECR higher than 12% (as- 
measured); 

• the same (approximately) tendency of the hydrogen absorption at 900 C as at 1100 C leads to the same 
(approximately) mechanical properties of the oxidized cladding. 

Unfortunately, it is impossible to continue the appropriate comparison for the second part of experimental 

data (950 C and 1000 C) in connection with the limited data on the hydrogen content but nevertheless, the 


3.57 











































obtained results allow to assume that the transition to the accelerated hydrogen absorption will take place 

with approximately the same (or a little less) ECR as that for the temperature 1100 C. 

As for revealed contradictions, then the following may be referred to those: 

• according to the results of visual observations (see Fig. 3.55, Fig. 3.56) the claddings oxidized at 900- 
1000 C were to demonstrate the worst mechanical properties and, consequently, to show the maximum 
values for the hydrogen absorption. However, this is not so or not quite so as it can be seen from the 
above stated analysis; 

• according to the data characterizing the allotropic phase transformation in the Zr-l%Nb alloys (see Fig. 
3.54) the test results with the El 10 cladding oxidized at 800 C may be logically explained by the fact 
that the (3-phase fraction in the El 10 cladding at this temperature is very small and, consequently, the 
general tendency of the cladding behavior corresponds to those of the low temperature oxidation of the 
E110 alloy in the area of the a-phase existence. All previous investigations have shown that the El 10 al¬ 
loy has the corrosion resistance in the a-phase temperature range; 

• but in accordance with the same data (see Fig. 3.54), it is quite difficult to explain by what the oxidation 
conditions in the temperature range 900-1100 C differ (even if we take into account the effects associ¬ 
ated with heating rates) as it is evident that the higher the temperature, the closer the Zr matrix composi¬ 
tion approaches the P-phase that must provide the avoidance of the breakaway oxidation. 

In connection with the list of revealed contradictions, it is appropriate to note the following: 

1. The availability of such contradictions as a rule indicates that the number of key factors taken into ac¬ 
count on analyzing is less than it is necessary. 

2. The following factors may be referred to the unaccounted ones: 

• the behavior of alloying elements (these aspects of the problem will be considered in the section 3.3.5 of 
the report; 

• the behavior of oxygen and a-Zr(O) phase as a function of temperature and the size of Zr-matrix grains 
as a function of temperature. 

• As for the behavior of oxygen a-Zr(O) phase and Zr-matrix as a function of oxidation temperature the 
appropriate data is organized in Fig. 3.60. 







ZrO, 


950 C 
#141-5 


800 C 


ECR=11-12.3% 


prior p 


100 Lilli 


Fig. 3.60. The comparison of the El 10 microstructure after a double-sided oxidation 

at different temperatures 


3.58 
























































The analysis of these data shown that: 

• the thickness ot a-Zr(O) phase at 800-950 C is significantly lower than that at 1000-1100 C in spite of 
difference in the oxidation level: 11-12.3% ECR at 800-950 C and 7.7-8.9% at 1000-1100 C; this effect 
leads to the fact that the thickness of the prior P-phase is approximately the same for both of the 
considered sets of data. But it is obvious that the prior P-phase thickness is one of most important factors 
that determine the residual ductility margin; 

• besides, the size of Zr-matrix grain at the lower temperature is a little less than that at 1000-1100 C; 

• moreover, the comparison of these two sets of experimental data shows that the tendency to the 
tormation of a-Zr(O) needles that penetrate into the prior P-phase and the tendency to the formation of 
a-Zr(O) layers along large grains of the prior P-phase are tvpical only for the temperature range 1000- 
1100 C. 


The revealed features of the El 10 behavior at different temperatures allow to understand that multiparametric 
elfects determine the cladding oxidation behavior under these conditions. But the resulting component of the 
combined effects from many factors may be characterized in the following way: 

• the oxidation at 800 C is the upper threshold for a good behavior of the El 10 cladding providing the 
approximate accordance between the experimental zero ductility threshold and the safety criterion; 

• the oxidation at 900-1100 C results in the fact that the experimental zero ductility threshold is lower than 
the safety criterion. 


The last position in this cycle of works was concerned with the investigation of peculiar features of the El 10 
cladding oxidation and embrittlement at the temperature 1200 C. The comparative data base characterizing 
this item of research is presented in Fig. 3.61. 


O'' 


O 

T3 


vi 

o 

a: 


70 


60 


— 50 


h= 40 


2 30 


20 


10 


0 


8 








o 

1200 C 

1100C 


1 

1 

1 






1 

1 

1 

1 













1 

1 

1 

-PA-- 


12 


16 


Measured ECR (%) 


1200 


£ 

c. 

c. 


c 

o 


O 

o 


800 


400 


0 




o 

o 

1100 N 

1200 N 

i 

i 

i 

i 





i 

i 

J 

i 

i 

i 

i 

i 

i 

1 o 



i 

i 

n 

1 

1 

1 

1 

1 

1 

II 

hi*noo 


0 


8 

ECR (%) 


12 


16 


Fig. 3.61. The characterization of the El 10 cladding behavior after a double-sided oxidation at 1200 C 


The analysis of obtained data leads to the following conclusions: 

• the zero ductility threshold of the El 10 cladding oxidized at 1200 C is most likely not better than that at 
1100 C; 

• the tendency to the decrease of hydrogen absorption was observed in accordance with the results of this 
test. But it is known that the tendency to the increase of the oxygen content in the p-phase characterizes 
the cladding behavior at this temperature in the comparison with 1100 C. 


3.59 
































• the E110 claddings continue to demonstrate the tendency towards the breakaway oxidation under this 
temperature. 

3.3.5. The sensitivity of the behavior of Russian niobium-bearing alloys 
to the alloying composition 

In the previous section of the report it was noted that: 

• the oxidation behavior of the Russian Zr-l%Nb cladding (El 10) and French Zr-l%Nb cladding (M5) is 
different; 

• the analysis of appropriate effects has shown that the first difference that draws the attention is the 
following: 

- the initial oxygen concentration in the El 10 cladding is 400 ppm (for claddings used in these tests); 

- the nominal oxygen concentration in the M5 cladding is 1350 ppm [35]; 

• the oxidation behavior of zirconium based claddings is sensitive to the concentration of such elements as 
Sn and Fe in the alloy. 

To verify the sensitivity of the niobium-bearing cladding behavior to these factors, several tests were 

performed with two following cladding types: 

• the Ell OK cladding: this is the El 10 cladding with the increased oxygen concentration (up to 
1100 ppm); 

• the E635 cladding: this is the niobium-bearing cladding with the following alloying composition: (Zr- 
l%Nb-1.2%Sn-0.35%Fe [41], 

The summary of test results with the El 10K cladding is presented in Fig. 3.62. 


3.60 



Fig. 3.62. The summary of results characterizing the E110K behavior under oxidation 

and ring compression test conditions 


The analysis of obtained results allows to establish with confidence the following: 

• the oxygen concentration increase in the El 10 alloy does not lead to the elimination of the effect of of 
the early breakaway oxidation; 

• the zero ductility threshold for the El 10K oxidized cladding is not higher than that for the El 10 cladding 
with the standard oxygen concentration. The E110K cladding has the same tendency of the hydrogen 
absorption by the prior (3-phase. 

Thus, these results bring to the conclusion that the general difference in the El 10 and M5 cladding behavior 

is not connected with the initial oxygen concentration in the cladding material. The cladding response on the 

variation of Sn and Fe composition in the cladding material was studied in the experiments with the E635 

cladding. The data base characterizing this test direction is presented in Appendixes B and F of the report. 

The organized results of these tests are shown in Fig. 3.63. 

The consideration of test results has led to the following conclusions: 

• the oxidation at 1000 C leads to more unfavorable consequences than those occurring at 1100 C, clear 
indications of the breakaway oxidation are observed on the appearance and microstructure of the 
cladding oxidized at 1000 C; 

• the appearance and microstructure of the E635 cladding oxidized at 1100 C are noticeably better than 
those of the El 10 cladding at 9.3% ECR: 

• the hydrogen measurements confirm these observations: 

- the hydrogen content in the E635 cladding oxidized up to the 9.3% ECR at 1100 C is significantly less 
than that in the El 10 cladding; 


3.61 














































- the hydrogen content in the E635 cladding oxidized at 1000 C even higher than that in the El 10 
cladding at the 5.3% ECR; 

- the results of ring compression tests performed with the E635 cladding oxidized at 1000 C are in a good 
agreement with the above listed observations. In accordance with these results, all general characteristics 
for the E635 cladding oxidized at 1000 C are not better than those for the El 10 cladding; 

- the results of mechanical tests with the E635 cladding oxidized at 1100C are rather contradictory as 
some rings have demonstrated the improvement of the mechanical behavior and the increase of residual 
ductility margin, but many rings have demonstrated the same results as those obtained for the El 10 
cladding or even worse ones. 

The results of this analysis show that apparently, the chemical composition of niobium-bearing alloys is the 
important factor for the behavior of these alloys under high temperature oxidation conditions. However, tests 
of the E635 cladding for just another time bring us to the conclusion that there are some other factors unac¬ 
counted in the present analysis that influence the oxidation mechanism. The analysis of these will be contin¬ 
ued in 4 chapter of the report. 

3.3.6. Interrelation between the zero ductility threshold and the temperature 
of mechanical tests 

All test results presented in the previous report sections were obtained in mechanical tests at the temperature 
20 C (room temperature). However, it is evident that for safety practical analysis, it is of importance to un¬ 
derstand to what extent the zero ductility threshold is sensitive to the temperature at the stage of proceeding 
from quenching during reflood to post-LOCA cooling. It is known that this stage is characterized by the 
achievement of the saturation temperature during reflood. This temperature may slightly differ for different 
reactors, however, by the existing tradition, the value of that is estimated as 135 C. 

The preliminary analysis of results from the studies performed earlier allowed to establish the following: 

• there are very limited data characterizing the oxygen induced embrittlement in this temperature range 
(20-135 C); 

• quite a number of investigations were performed to study the effect of the hydrogen induced 
embrittlement. 


3.62 


Middle oxidation Low oxidation 


1100C 


ECR=6.8% 



#135 


1000 C 


ECR=5.3% 


#138 



1100C 


ECR=9.3% 



#134 


1000 C 


ECR=9.4% 



#127 


1000 C (9.4% ECR) 


Etched , #138-5 



1100 C (9.3% ECR) 


Etched , #135-5 



80 


60 


o 


= 40 


■y. 
fl J 


20 


1200 


-E110, 1100 C 

a E635, 1100 C 
E635, 1000C 


£- 800 


\ A 

1 

1 

\ 


CC 

O 


o 

CD 


400 


1 

i 


A 

A 


-I— 

6 


10 


e-635-ull-m-v3 

- 1 

14 


□ 


E110. 1100C 
E635, 1100C 
E635, 1000C 


O A | 

l 


Measured ECR (%) 


4 8 12 

Measured ECR (%) 


16 


Fig. 3.63. The characterization of the E635 cladding behavior under oxidation 

and mechanical test conditions 


3.63 
























































So, as for the oxygen embrittlement, the most part of published data was devoted to the studies of the 
zirconium cladding mechanical behavior as a function of the initial oxygen concentration in the cladding 
material. In accordance with Russian data (for Zr-l%Nb) and other data for alloys of the zircaloy type, the 
elongation of the claddings with oxygen concentration 0.05-0.16% (by weight) in the a-phase is not 
sensitive to the temperature within the range 20-200 C. The previous Russian investigations performed in the 
Bochvar Institute (VNIINM, V.Tonkov) have shown that the ductility of the El 10 oxidized claddings as a 
function of temperature of mechanical tests in the range 20-200 C is subjected to the following regularities: 

• Zr0 2 and a-Zr(O) ductility is not increased in this temperature range; 

• a-phase (prior P-phase) ductility is the function of the temperature especially in the range 100-200 C; 
however, the researcher associated this effect with the hydrogen behavior in the cladding material. 

Unfortunately, this issue (the oxygen induced embrittlement) did not meet with due elucidation even in such 
a complete monograph which is the monograph of D.L. Douglass [42]. Though, the review made by D.L. 
Douglass contains the illustration that demonstrates that the samples manufactured from Zr-2.5%Cu alloy 
and oxidized up to 0.6% of the oxygen content in the zirconium matrix elongate from 2 up to 31% on the 
temperature increase from 20 up to 200 C. 

The previous investigations performed by the authors of this report with unirradiated and irradiated Zr-l%Nb 
claddings have shown that the cladding elongation is not sensitive to the temperature of mechanical tests 
(20-200 C) in the presence of irradiation damage effects (50 MWd/kg U), low oxidation (oxide thickness 
5 pm) and low hydrogen concentration (30 ppm) [19]. 

As for the hydrogen embrittlement, the appropriate effects revealed in a number of studies may be 
characterized in the following way: 

• the oxidized cladding fracture resistance is a strong function of precipitated hydrides; 

• the cladding ductility margin is a function of hydride concentration, size, orientation and morphology; 

• the ductility of hydrides is a function of the temperature in the range from approximately 100 C and 
higher; 

• it was experimentally demonstrated that brittle cracks initiated in the brittle a-Zr(O) layer slow down and 
are not developed in the prior P-phase while the hydride ductility is increased. In its turn, this effect 
results in the fact that the cladding elongation is increased also. 

Besides, some additional observation of importance should be noted: 

• the hydride-related embrittlement of such alloy as the irradiated Zry-4 is a function of a transformed beta 
microstructure [43]; 

• the temperature of ductile-brittle transition (DBT) is slowly increased with the hydrogen concentration 
increase [42]; 

• for the low hydrogen concentration in the prior P-phase, the transition from the brittle to ductile fracture 
is very sensitive to the temperature [42]; 

• the cladding ductility is not sensitive to the hydrogen concentration up to 70 ppm [44]; 

• the unirradiated cladding ductility at the room temperature is decreased very sharply down to low values 
with the hydrogen content of about 700 ppm [45]; 

• the worst ductile behavior was demonstrated at the room temperature in the cladding material with 
uniformly distributed hydrides [46], 

The obvious illustration for some of the listed effects is demonstrated in Fig. 3.64 [47], In accordance with 
these data, the cladding samples with lower hydrogen concentration (180 ppm) have demonstrated 
significantly higher sensitivity to the temperature of mechanical tests in the range 50-150 C than that of the 
cladding samples with hydrogen concentration of about 700 ppm. At 200 C, the mechanical behavior of both 
types of samples was quite ductile. 


3.64 




700ppm HYDROGEN 




■> Displacement 


Fig. 3.64. Load-displacement diagrams for two Zry-2 samples hydrided (specially) 
up to 180 ppm and 700 ppm, respectively, as a function of 
the temperature mechanical tests of the bending type (reprinted from [47]) 


To reveal the temperature effects in the El 10 oxidized and hydrided cladding after the oxidation tests 
performed in the frame of this work, the following approach w as applied: 

• the representative scope of mechanical tests (ring compression tests) was performed at 135 C with the 
E110 cladding samples oxidized at 800-1200 C and F/F, F/Q combinations of heating and cooling rates; 

• several reference ring compression and ring tensile tests were performed with the El 10 oxidized 
cladding samples in the temperature range 20-300 C. 



Measured ECR (%) 

Fig. 3.65. Dependence of the El 10 cladding ductility on the ECR (oxidation at 1100 C) 

and temperature of ring tensile tests 


In accordance with these results the following comments can be made: 

• the ductile cladding sample (#47, 7% ECR. the residual ductility 55-65% according to ring compression 
tests, low hydrogen content 30 ppm) demonstrated maximum sensitivity to the temperature of ring tensile 
tests in the range 20-135 C. The total elongation of this sample increased from 7% at room temperature 
to 17.5% at 135 C; 

• the almost brittle sample (#49, 7.5 % ECR. the residual ductility 1.3-3.6 % according to ring compres¬ 
sion tests, critical hydrogen content ~ 500 ppm) demonstrated: 

- pronounced sensitivity to the temperature of ring tensile tests in the range 20-135 C. The total elonga¬ 
tion increased from 1.5 to 5% respectively; 

- very high sensitivity to the tensile test temperature in the range 135-200 C. Total elongation of this 
quite brittle sample increased from 5 to 16%; 


3.65 

























- further increase of tensile test temperature to 300 C did not lead to the increase of ductility; 

• the brittle sample with very high hydrogen content in the prior (3-phase (#25, 9.9% ECR, hydrogen con¬ 
tent 1000 ppm ) shown a low sensitivity to the tensile test temperature between 20 and 135 C and high 
sensitivity in the range 135-300 C, where the total elongation increased from 1.7 to 10.6%. 

Before we analyze the above stated comments, it will be useful to supplement this data base with results of 


the ring compression test performed at 20-300 C. The results are presented in Fig. 3.66. 



Fig. 3.66. Residual ductility of two El 10 samples oxidized at 10 and 11.7% ECR (1100 C) 

as a function of temperature ring compression tests 


The obtained data lead to the following observations: 

• the sample (#68), which was oxidized somewhat higher than the zero ductility threshold (10% ECR) up 
to the hydrogen concentration of about 1000 ppm, has demonstrated definite sensitivity to the 
temperature of mechanical tests ranging from 20 up to 135 C; the residual ductility of this sample 
significantly increase in the range 135-200 C. However, we did not manage to reveal the influence of the 
temperature range 200-300 C over this sample ductility due to the fact that the maximal grip 
displacement (in these tests corresponded to the residual ductility of about 65%) was already achieved at 
the 200 C; 

• the brittle sample (#36, ECR=11.7%) with a very high concentration of hydrogen (1500 ppm) was 
insensitive to the temperature of mechanical tests in the range 20-135 C, this sample ductility sharply 
increased in the temperature range 135-200 C, after that, the additional ductility increase was observed 
in the temperature range 200-300 C. 

If we summarize the results of the preliminary analysis then the following tendencies may be noted: 

• the less is the hydrogen concentration, the higher sensitivity to the temperature increase is demonstrated 
by El 10 oxidized samples in the temperature range 20-135 C; 

• the influence of the temperature of mechanical tests within the range 20-135 C over ductility of the El 10 
oxidized samples manifests itself still at the hydrogen concentration 1000 ppm, however, on the further 
hydrogen concentration increase (up to 1500 ppm), the cladding ductility is insensitive to the temperature 
within this range; 

• the temperature increase up to 200 C leads to the significant increase of the cladding ductility in the 
whole studied range of the hydrogen concentrations; 

• the temperature increase up to 300 C will influence the cladding ductility the more noticeably, the higher 
it was hydrided at 20 C. 


3.66 








































The peculiar features revealed tor the mechanical behavior of the El 10 oxidized cladding may be clarified 
still more with the help ot the representative data base obtained on the basis of ring compression tests per¬ 
formed at 20 C and 135 C. The first organized data characterize the relationship between the residual ductil¬ 
ity of the E110 oxidized cladding and the hydrogen concentration at two temperature levels (see Fig. 3.67). 



0 400 800 1200 1600 

EE content (ppm) 

Fig. 3.67. The sensitivity of the El 10 residual ductility' (800-1200 C, F/f and F/Q) 
to the hydrogen concentration at 20 and 135 C 

These data allow to make the following general conclusions: 

• the residual ductility of the El 10 oxidized cladding is a function of the hydrogen concentration; 

• the El 10 cladding has a very high level of residual ductility (higher than 50%) at 20 C within the interval 
of hydrogen solubility at 20 C (0-100 ppm); 

• the residual ductility at 20 C decreases vary sharply in the narrow range of the hydrogen concentration: 
100-200 ppm: 

• the zero ductility threshold at 20 C is situated between 400-800 ppm of the hydrogen concentration; 

• taking into account that the hydrogen solubility limit and ductility of hydrides are increased at 135 C. the 
E110 cladding samples demonstrate a very high level of residual ductility in the range of the hydrogen 
concentration 0-500 ppm; 

• at 135 C, the decrease of residual ductility takes place very sharply with the hydrogen concentration 
500-700 ppm and the zero ductility threshold corresponds to 900 ppm of the hydrogen concentration at 
this temperature. 

The relationship between revealed effects and the oxidation level may be characterized using the second 
organized data presented in Fig. 3.68. In this case, the increment of residual ductility at 135 C (residual duc¬ 
tility at 135 C minus residual ductility at 20 C) was determined as a function of the ECR. 

As it was already mentioned above, the results of ring compression tests do not allow to observe the effect of 
ductility increase with the temperature increase in mechanical tests for samples with high residual ductility at 
20 C because the grip displacement of test machine was specially limited. In this connection, the increment 
of residual ductility at 135 C was not revealed in the range 0-6.5% (0-100 ppm of the hydrogen 
concentration). The ECR range of 6.5-8.3% characterizes a sharp decrease of residual ductility at 20 C, the 
sharp increase of the hydrogen content up to 700 ppm and a fast increase of increment of residual ductility at 


3.67 





















135 C up to the maximum value that corresponds to the zero ductility threshold at 20 C (the critical ECR is 
8.3%). After that, the increment of residual ductility at 135 C decreases fast down to zero that is associated 
with the increase of oxygen and hydrogen concentration in the prior P-phase. 



Fig. 3.68. The data characterizing the sensitivity of the El 10 residual ductility at 135 C to the ECR 

(900-1100 C, F/F and F/Q) 

Thus, this cycle of investigations allowed: 

• to reveal the major effects associated with the hydrides behavior as a function of temperature; 

• to reveal the sensitivity of the El 10 residual ductility to the temperature. 

But the analysis shows that one aspect of this issue remained not quite clear, namely, the consideration of 
temperature effects, associated not so much with the hydrogen embrittlement as with the oxygen embrittle¬ 
ment. Therefore, the analysis will be continued in chapter 4 of the report. 

3.3.7. The analysis of representativity of the zero ductility threshold deter¬ 
mined due to ring compression tests 

It is obvious that the oxidized claddings of fuel rods will experience complicated multi-dimensional loadings 
during the real LOCA. Therefore, the problem of representativity of the zero ductility threshold determined 
using relatively simple ring compression tests is discussed simultaneously with performing studies of this 
type and during the time of this performance. The cycle of appropriate investigations performed in the frame 
of this work is the immediate contribution into the solution of this problem. 

The approach, developed to obtain the comparative data for this issue analysis included the following items: 

• the performance of several tests with other types of mechanical loading for the cladding samples and the 
comparison of the whole set of obtained results; 

• the extension of the test data base involved into the determination of the zero ductility threshold due to 
the processing of data characterizing the maximum loads at the fracture; 

• the comparison of macroscopic and microscopic data to confirm brittle or ductile fracture of oxidized 
claddings. 

The test performed in accordance with the first item of this list included the ring tensile tests and three-point 
bending tests. The procedures for these tests are described in sections 3.2.2.2, 3.2.2.3 of the report. The com¬ 
parative data characterizing results of ring tensile and ring compression tests are presented in Fig. 3.69. 


3.68 
























20 


16 


12 


SO 

c 

o 


- 8 


o 


Ring tensile tests 















1- 

\ j 

\ 


TE ECR 20-c grf 

— 

* 

- * 

— 


8 

Measured ECR (%) 


10 


12 


80 


60 


£. 


= 40 


20 


Ring compression tests 


— 

\ 

\ 





\ 

\ 

\ 





\ 

\ 

\ Ze 

\ 

ro dnailin threshold 

A ' ] 


e! IOk-5-e frf 

\ 

\ 

4 

ri- 




8 10 
Measured ECR (%) 


12 


14 


Fig. 3.69. The comparison of zero ductility thresholds determined from the results of ring tensile and 

ring compression tests (El 10, 1100 C) 


The obtained data indicate clearly that there is no difference in the zero ductility threshold determined using 
the ring tensile tests or ring compression tests. The results of three-point bending tests presented in Fig. 3.70 
allow to conclude that the zero ductility threshold determined due to this type of tests is higher than that de¬ 
termined using the ring compression tests (11.8% and 8.3% ECR. respectively). 



Fig. 3.70. The zero ductility threshold of the El 10 cladding determined 
due to three-point bending tests (1100 C, F/F) 

Thus, the comparative data base obtained due to different types of mechanical tests shows that ring 
compression tests allow to obtain the conservative estimation of the zero ductility threshold. This conclusion 
made on the basis of macroscopic tests is confirmed by the results of microscopic observations obtained 
using the fractography examinations. The fragments of two rings were selected for the fractography 
examinations after the fracture under the ring compression test conditions: 

• a brittle ring: 1100 C, 8.2% ECR, double-sided oxidation; 

• a ductile ring: 1100 C, 6% ECR, single-sided oxidation. 

Two fracture surfaces of the brittle ring fragments were studied in detail (see Fig. 3.71): 

• fracture surface characterizing the behavior of sample segment, which experienced the compression 
stresses on the outer surface and tensile stresses on the inner surface (the first fracture surface); 


3.69 


































• fracture surface characterizing the behavior of other part of cladding, which experienced the tensile 
stresses on outer surface and compression stresses on the inner surface (the second fracture surface). 



Fig. 3.71. SEM micrographs for fracture surfaces of the El 10 brittle cladding 


The examinations of the first fracture surface performed using the SEM micrographs with a high magnifica¬ 
tion (see Fig. 3.72) have shown that: 

• the oxide layers on the outer and inner cladding surface are of the columnar structure and the oxide 
thickness is 15 pm (Fig. 3.72a,b); 

• a typical pattern for a-Zr(O) surface is of the cleavage type (see Fig. 3.72b); 

• the fracture pattern of the prior P-phase layer is quasi-cleavage (Fig. 3.72c), the fracture surface may be 
characterized as the “terrace” type (Fig. 3.72d), separate small regions of a dimple rupture are observed 
on the boundary between the a-Zr(O) and prior P-phase layers (Fig. 3.72e); besides, the transition from 
the quasi-cleavage fracture to the cleavage fracture is revealed in this region. 

The structure of the second fracture surface as a whole does not differ from the first fracture surface though, 
a somewhat higher number of ductile fracture regions was fixed in this sample. Moreover, the fracture sur¬ 
face of the prior P-phase in this sample is characterized by the mixed type of the fracture pattern: the combi¬ 
nation of quasi-cleavage facets and the dimples of the ductile fracture (Fig. 3.72f). 

The analysis of the fracture surface in the ductile cladding sample confirms that the surface pattern of the 
prior P-phase is typical for the ductile fracture (see Fig. 3.73). Only separate small regions of the quasi¬ 
cleavage fracture type were observed in this sample. Thus, the fractography data have shown that small areas 
with the residual ductility are present in the material of even brittle samples but on the whole, the reasonable 
agreement between the microscopic and macroscopic assessments of the zero ductility threshold is observed 
in the fractography examinations. 


3.70 








Fig. 3.72. High magnification SEM micrographs of fracture surface regions 

of the E110 brittle cladding 



Fig. 3.73. The SEM micrograph for the fracture surface of the El 10 ductile sample 

It should be noted that the performed cycle of investigations employing ring compression tests for the 
determination of the zero ductility threshold in the El 10 oxidized cladding allowed to reveal some more 


3.71 




























issues connected with the representativity of these tests. The following two of those arc considered in this 
report: 

• the analysis of correlation between the fracture load and fracture displacement; 

• the additional analysis of the sensitivity of the ring compression test results to the test procedures. 

The first issue essence my be characterized in the following way: from the practical point of view, the sig¬ 
nificance of what strain the oxidized cladding has undergone during post-quench actions is not so important 
as that of what maximum load the cladding can stand before its fragmentation. It may be assumed that this 
was the reason for the authors of one of the recent papers dedicated to the M5 cladding behavior under the 
LOCA relevant conditions to present the characterization of the M5 fracture on the basis of the analysis of 
maximum loads at ring compression tests [46]. The approach of this type was the subject of the analysis per¬ 
formed in the frame of this work also. The organized data to clarify the appropriate issue are presented in 
Fig. 3.74. 

1000 


800 

Z 

a 

o 

- 600 
c3 

B 

• 

x 

03 


400 


200 

0 10 20 30 40 50 60 70 

Residual ductility (%) 

Fig. 3.74. The maximum load on the El 10 oxidized sample as a function of residual ductility 

The obtained data show the following: 

• the decrease in the cladding residual ductility down to several percent affects its capability to withstand 
the load and indicates low strain hardening of the El 10 oxidized cladding; 

• there is a clear correlation between the zero ductility threshold and the sharp reduction in the strength 
properties of the El 10 oxidized cladding (thus, the maximum load is decreased 3-fold as the ECR is 
increased from 8% up to 12%). 

In other words, these data confirm that the zero ductility threshold determined as the zero residual ductility 
margin strongly corresponds to the appropriate critical load. But it should be noted that the employment of 
the critical load for the evaluation of the zero ductility threshold is not convenient as the critical load is not 
only the function of the ECR but it is also the function of the cladding sample length. Taking into account 
that there is no standard sample length for the ring compression tests and that different laboratories use 
samples with different lengths, in this case, it is impossible to compare the results. 

The second issue that will be touched upon in this section of the report concerns just the comparison of the 
E110 zero ductility thresholds obtained in the ring compression tests performed in different laboratories. The 
analysis of comparative data base presented in Fig. 3.75 is devoted to the consideration of possible reasons 
for differences in evaluations of the El 10 zero ductility threshold obtained by different laboratories. 

The first observations concerning this comparative data base may be formulated in the following way: 


pw 





o 

°c 

c 

>i 

o 

o 

■8°° 

%--$ . 

O O 

— « 

—--- 

c 

.--- o 

o 

o 

° O .... 

-- o 

OO 

° o 

° 8 

o 


f%° o 

o 

o 

o 









1100 C 
double-sided 


> 

\ _ 


— 

— 







-1 


3.72 




















• the relative order of the El 10 regression correlations for test data obtained at VNIINM [15], KfKI [18, 
25], NC Rossendorf [17, 24] and Scoda-UJP [10, 26] corresponds exactly to the order of the Zry-4 re¬ 
gression for these laboratories (see Fig. 3.28); this fact leads to the conclusion that this order is the func¬ 
tion of experimental procedures used in each laboratory; 

• in contrast to the data characterizing the Zry-4 mechanical behavior, the systematic difference is 
observed between the RRC KI/RIAR data and all other results; 

• the zero ductility threshold of the El 10 oxidized cladding is about 4.5-6% ECR in accordance with the 
VNIINM, KFKI, NC Rossendorf, Scoda-UJP results, while that is about 8.3% ECR in accordance with 
RRC KI/RIAR results. 



Fig. 3.75. The comparative data characterizing the El 10 residual ductility as a function of the ECR 
obtained on the processing of test data of different laboratories 

The special analysis results show that the following effects may be responsible for this difference: 

• systematic errors in the procedure of the weight gain determination; 

• differences in the El 10 cladding material used for the oxidation tests; 

• differences in the procedures of the oxidation tests (coolant type (water steam, water steam/argon 
mixture), heating and cooling rates) and differences in the procedures for the processing of load- 
displacement diagrams. 

As for the procedure of the weight gain determination, the data base presented in Fig. 3.29 may be employed 
to assess the scale of this effect. These data show that the KFKI data really somewhat underestimate the 
weight gain in the Zry 4 cladding in comparison to the RRC KI/RIAR and NC Rossendorf data, but this 
effect is very small; the data presented in Fig. 3.28 allow to assume that the weight gain was very 
underestimated in the VNIINM tests. Besides, it is known that the Scoda-UJP tests were performed with the 
E110 claddings manufactured using the previous method of the El 10 alloy fabrication, namely, the El 10 
claddings were fabricated from the iodide zirconium in contrast to iodide/electrolytic Zr used to manufacture 
the El 10 cladding employed in the RRC KI/RIAR tests. The nature of the El 10 cladding material used in the 
VNIINM, NC Rossendorf and KFKI tests is unknown. Besides the steam/argon mixture was used in the 
some of these tests. The sensitivity of test results to this parameter is not quite understood also. 

And, finally, one more important potential cause for revealed differences must be considered. This cause is 
associated with the procedure for the preparation of the cladding samples for the mechanical tests. So, in 
section 3.2.2 it was demonstrated that results of the ring compression tests are not the function of the sample 
length. But this conclusion is referred to the procedure adopted in this study according to which the end parts 


3.73 
















of the oxidized cladding were cut off and the mechanical tests with these parts of the cladding were not 
performed in contrast to the KFKI and VNIINM tests. Besides, the VNIINM tests were performed with very 
long oxidized samples (30 mm). The end parts of 20 mm samples were apparently removed in the NC 
Rossendorf tests because in accordance with the description of tests, two ring samples 8 mm and 5 mm long, 
respectively, were prepared from each of 20 mm oxidized claddings for compression tests and metallography 
investigations. But in our opinion, the lengths of cut off ends may appear to be not enough to compensate the 
effects described below. As for the Scoda-UJP tests the 7 mm rings were prepared from 30 mm oxidized 
samples. 

The effects of the end parts of the El 10 oxidized cladding may be characterized in the following way: 

• the cladding oxidation takes place not only on the sample side surfaces but also on end surfaces of that. 
This results in the fact that the prior p-phase small cladding sample (for illustration, the sample 6 mm 
long was used in the KFKI tests) absorbs the oxygen from four sides. It is evident that the residual 
ductility of such sample will be lower than that of the sample oxidized from two sides only. In this case, 
this effect will be especially expressed at relatively low ECRs; 

• the fact that the similar process taking place at the hydrogen absorption is still of more importance as the 
E110 alloy embrittlement occurs not so much according to the mechanism of the oxygen embrittlement 
as to that of the hydrogen embrittlement. 

The effect of the hydrogen absorption by the end surface may be qualitatively illustrated by the results pre¬ 
sented in Fig. 3.76 (Reprinted from Reference 48, Fig 5b). 



Fig. 3.76. Hydrogen distribution along the width of a special Zry-4 sample 

Special investigations performed in JAERI [48] allowed to see that the end effect is strongly expressed on the 
length of 3 mm under these test conditions. The hydrogen concentration is 2.5 times reduced along the whole 
length. 

Besides, it is necessary to point out once more that the hydrogen absorption by the El 10 cladding is the con¬ 
sequence of the breakaway effect. In this case, as it was repeatedly noted earlier, the precise connection is 
between stresses in the oxide layer and the initiation of this effect. A special analysis performed during this 
research allowed to establish that the initiation of the breakaway effect takes place at the end parts of the 
cladding sample as, apparently, this part of the cladding sample represents a special concentrator of stresses 
occurring on the boundary between the end and side cladding surfaces. This statement is obviously demon¬ 
strated by the data presented in Fig. 3.77. 

The obtained data obviously indicate that the initiation of the breakaway oxidation and, consequently, the 
initiation of the cladding hydriding, and then, the cladding embrittlement takes place in the end part of the 
oxidized cladding noticeably earlier than in the sample basic part. 

Taking into account this analysis results, it may be assumed that the difference between the RRC KI/RIAR 
assessment of the El 10 zero ductility threshold and assessments of this threshold performed in other labora¬ 
tories is explained in the first turn by the fact that the RRC KI/RIAR test data are independent of the end 


3.74 






effects while the results of other investigators are overburdened with the end effect. Though, it is evident that 
other tactors listed in the discussion of this issue additionally contributed into revealed differences. 


#63 

#111 

#46 

6.1% ECR 

8.5% ECR 

6.5% ECR 



Fig. 3.77. Demonstration of the end effects on the El 10 oxidized cladding samples 


3.3.8. Consideration of the zero ductility threshold of the El 10 cladding as a 
function of the irradiation effect 

It is important to note that this stage of research does not pretend to be completed. The goal of this stage was 
to obtain the first experimental data characterizing the scale of the irradiation effect as applied to the El 10 
cladding. To study the appropriate phenomena, eleven irradiated claddings refabricated from the VVER high 
bumup fuel rods (50 MWd/kg U) were tested. The initial characteristics of irradiated claddings are presented 
in Appendix A-3. The oxidation and hydrogenation of the El 10 irradiated cladding were characterized by 
the following values before the oxidation tests: 

• the outer oxide thickness is 5 pm; 

• the inner oxide thickness is 0 pm; 

• hydrogen content is 47 ppm. 

The results of the irradiated cladding tests are presented in Appendixes B and I of the report. These tests 
consisted of two stages: 

1. Scoping oxidation tests performed at the test modes with S/S (slow/slow) and S/F (slow/fast) combina¬ 
tions of heating/cooling rates. 

2. Basic oxidation tests performed at the F/F combination of heating and cooling rates with the variation of 
temperature in the range 1000-1200 C. 

The analysis of the cladding appearance and microstructure after the basic oxidation tests at 1100 C allows to 
note the following (see Fig. 3.78): 


3.75 

















any indications of the breakaway oxidation are not observed on the outer cladding surface up to 7.0% 
ECR; 


• insignificant indications of the breakaway oxidation occur at the higher ECR, however, these indications 
are expressed significantly weaker than those for the unirradiated cladding; 

• the oxidation behavior of the irradiated cladding differs considerably by the following processes: 

- the obviously expressed tendency towards the increase of the oxide thickness on the inner cladding 
surface in comparison with the outer oxide thickness; 

- the obviously expressed tendency towards the formation of the oxide lamination and the tendency 
towards the oxide spallation starting from 7.7% ECR on the inner cladding surface; 

- the more expressed tendency towards the generation of the a-Zr(O) phase along the grain boundaries of 

the prior [3-phase. _ 


before the test 

ECR-0.5% 



after the test (#14) 



ECR=8.3 % MBHMBBMB 




Sample 

Before oxida¬ 
tion tests 
(Polished) 


#20-4 

ECR=6.3 % 
(Etched) 


#10-4 

ECR-7.7 % 
(Polished) 


#14-4 

ECR-8.3 % 
(Polished) 


Outer surface 


Inner surface 






Fig. 3.78. The appearance and microstructure of the El 10 irradiated cladding before the tests and 
after the oxidation tests at 1100 C and F/F combination of heating and cooling rates 


3.76 


























































It may be assumed that the revealed peculiar features are associated, first, with the presence of fission 
products on the inner cladding surface and with the participation of some of those in the oxidation reaction 
and, second, with the change of the cladding microstructure during the base irradiation. 

As for the dependence of the oxidation behavior on the test conditions, then the performed studies have 
shown that: 


• the irradiated cladding oxidation under slow transient conditions (S/S combination of heating and cool¬ 
ing rates) leads to the macroscopic effects of the oxide spallation at 10.5% ECR (see Fig. 3.79); 




the oxidation at 1200 C leads to the significant decrease or disappearance of the breakaway effect from 
the microscopic point of view, nevertheless the microstructure demonstrates the "hydrogen-modified" 
type (see Fig. 3.79). 


1100 C #2 



Etched 


Polished 


#18-4 

ECR=16.0% 

T=1200C 



100 fim 






Fig. 3.79. The appearance and microstructure of the El 10 oxidized cladding after the slow transient 

oxidation at 1100 C and standard oxidation at 1200 C 


The ring compression tests performed with the El 10 oxidized irradiated claddings allows to obtain the results 
presented in Fig. 3.80. The preliminary analysis of these results shows that a general tendency towards the 
decrease in the zero ductility threshold (ZDT) is observed in the irradiated claddings. To extend the data base 
for the more accurate analysis of specific physical processes in these claddings, the following additional in¬ 
vestigations were performed: 

• the microhardness measurement; 

• the hydrogen content measurement. 


3.77 





















Fig. 3.80. The residual ductility of the El 10 irradiated cladding as a function of the ECR 

The comparative data presented in Fig. 3.81 to characterize the microhardness of different types of the El 10 

cladding (unirradiated, irradiated, oxidized and unoxidized) allow to make the following important 

observations: 

• the initial microhardness of the El 10 irradiated cladding (before the oxidation) is higher than that of the 
E110 unirradiated cladding by 60 kg/mirf; 

• the microhardness of the El 10 unirradiated oxidized cladding (7.0% ECR) with a high margin of 
residual ductility (sample #47) is about 250 kg/mirf and (as it was demonstrated earlier, see Fig. 3.44) 
the microhardness of the El 10 unirradiated oxidized cladding (ECR=8.2%) at the zero ductility threshold 
(sample #41) is the same; 

• the microhardness of the El 10 irradiated oxidized cladding (6.5% ECR) with a significant margin of 
residual ductility (sample #17) is practically the same as that in the El 10 irradiated oxidized cladding 
(7.7% ECR) close to the ZDT (sample #10); this microhardness is about 300 kg/mm 2 ; 

• the microhardness of the El 10 irradiated oxidized cladding increases significantly in the ECR range 7.7- 
8.3% ECR and after that the microhardness is not practically changed up to 16% ECR. 



Fig. 3.81. Comparative data characterizing the microhardness 
of the E110 oxidized irradiated claddings 


3.78 



























The analysis ol these observations leads to the following conclusions: 

• some initial embrittlement effect is observed even in such a “good” irradiated cladding as the El 10 
cladding; 

• the oxygen concentration exceeding the oxygen embrittlement threshold (0.6-0.9% by weight) of the 
prior p-phase is apparently achieved in the El 10 irradiated cladding at the ECR>8.3%; 

• the zero ductility embrittlement threshold of the El 10 irradiated oxidized cladding is apparently 
determined by the same combination of factors that was determined for the El 10 unirradiated cladding, 
namely: by the combined effect of the oxygen and hydrogen embrittlement of the oxidized cladding prior 
P-phase. 

The continuation of this analysis performed using the results of the hydrogen concentration measurements 

allowed to reveal the following additional peculiar features for the behavior of the El 10 irradiated cladding: 

• in spite of the fact that the oxide spallation was obviously demonstrated in the oxidation mode under 
slow transient conditions (S/S combination of heating and cooling rates), all three cladding samples 
tested at slow heating had a relatively low hydrogen concentration in comparison with the samples tested 
at fast heating; 

• the cladding samples tested under slow heating conditions demonstrated the tendency towards the 
increase of the zero ductility threshold in comparison with the samples oxidized at fast heating. 

Taking into account all above mentioned considerations, two correlations were developed to characterize 

residual ductility of the El 10 irradiated oxidized cladding as a function of the ECR and oxidation modes (see 
Fig. 3.82). 



Fig. 3.82. The zero ductility threshold of the El 10 irradiated cladding 

after different oxidation modes 


It should be also noted that to minimize a possible error in the assessment of the zero ductility threshold at 
the F/F oxidation mode, the load-displacement diagrams were additionally analyzed. As it can be observed in 
the data presented in Fig. 3.82, the cladding sample #21 was tested at the ECR that was somewhat higher 
than the zero ductility threshold. The quantitative analysis of this discrepancy performed using the analysis 
of other load-displacement diagrams has shown that the zero ductility threshold for this oxidation mode F/F 
may be evaluated with a good accuracy as 6.5% ECR. 

The organization of the data base characterizing the hydrogen content in the El 10 irradiated oxidized clad¬ 
ding as a function of the ECR was performed in accordance with the similar approach (see Fig. 3.83). 


3.79 



















Fig. 3.83. The El 10 hydrogen concentration as a function of irradiation and ECR, 
residual ductility of the El 10 irradiated cladding as a function of hydrogen concentration 


The regression dependence presented in Fig. 3.83 characterizing the hydrogen content in the El 10 irradiated 
cladding was developed for the F/F oxidation mode in the temperature range 1000-1100 C. The appropriate 
dependence at the temperature 1200 C will apparently demonstrate a smoother increase of the hydrogen 
content as a function of the ECR. The comparison of obtained data with results obtained for the El 10 
unirradiated cladding shows that the hydrogen absorption takes place in the irradiated cladding still more 
intensively. However, coming back to the results of the analysis performed using the metallographic views, it 
may be assumed that special processes occurring on the inner cladding surface are responsible for revealed 
differences. 

Additional data presented in Fig. 3.83 to describe the relationship between residual ductility of the El 10 
irradiated cladding and hydrogen content (developed with regard to all test modes and all temperatures) are 
in a good agreement with the similar data obtained for unirradiated claddings. Nevertheless, it is impossible 
to exclude the fact that a definite peculiarity in the behavior of the irradiated cladding during the oxidation 
may be caused by the changes in its microstructure that took place during the base irradiation. Thus, the SEM 
investigations performed with the El 10 irradiated cladding allowed to reveal the following tendencies: 

• the cladding material structure is characterized by the a-Zr phase containing the global P-Nb 
precipitates; 

• the niobium concentration in the matrix is practically equal to zero. 

Taking into account the considered earlier oxidation effects associated with niobium, it may be assumed that 
in this case, the change of the Zr-l%Nb material structure results in some change of its oxidation behavior. 

3.3.9. The analysis of the El 10 oxidation kinetics 

To develop the El 10 oxidation kinetics, the parabolic law was used: 

AW 2 = Kt, 

where AW - the weight gain (mg/cirr); 
t- time (s); 

K - rate constant (mg 2 /cm 4 s). 


3.80 


















































Besides, it was assumed that the rate constant may be described by Arrhenius relation: 


K = A exp 


0 


\ 


R T 


where A - 


empirical coefficient (mg" cm" s): 


Q- activation energy (J mol); 


R - gas constant (J mol K); 
T - temperature (K). 


The processing ot test data started from the determination of K using each measured combination of AW and 
t. The averaged value of AW obtained in the several ring samples cut off from one 100 mm cladding sample 
was employed tor this goal. These data were used to develop the regression dependence presented in Fig. 
3.84. It should be noted that the data obtained at the F F and F Q combinations of heating and cooling rates 
only were used for this procedure. 

The obtained correlation is valided for the following oxidation duration as a function of the oxidation 
temperature: 


• 1073 K -> 29000 s; 

• 1173 K —> 4800 s; 


• 1223 K —» 5000 s; 

• 1273 K—> 1800s; 


• 1373 K —> 1800 s; 

• 1473 K-> 400 s. 



Fig. 3.84. Determination of the rate constant at the oxidation of the El 10 unirradiated cladding 

in the temperature range of 1073-1473 K 

The sensitivity of the El 10 oxidation kinetics to the irradiation effect may be preliminary estimated using the 
data presented in Fig. 3.85. 


3.81 

















20 


16 


s 

o 

t? >2 


03 


JC 

SJj 


<D 

£ 


E1I0,1000C 
• irradiated 


unirradiated 


k W 00 -ir-rrc-v 2 


200 


400 

Time (s) 


600 


800 


20 


16 • - 


p 

o 

6 


5 Q 


12 


£ 8 
sp 

£ 





k-JI 00 -ir-rrc-v 2 


E110, 

• 

1100C 

rradiated 

jnirradiated 

_ 








• 









200 


400 

Time (s) 


600 


800 


P 

so 

£ 

p 

5 

50 


20 


16 


12 


■H 8 

.50 

o 

£ 





kl 200 irrrc-v 2 


• 







jr 







E110,1200C 
• irradiated 

unirradiated 


f 




J_1 



10 


P 

u 

r-4 


0.1 


0.01 


0.001 


0 0001 




o un 

• irr 

irradiated 

adiatcd 

irradiated 


• 

© 

| 

- un 





1 « 

9 d- 

8 

O 




o 

o \ 





k-l T-rrc-2-v2 


200 400 

_ Time (s) 


600 


800 


0.0006 


0.0007 


0.0008 

1/T (K') 


0.0009 


0.001 


Fig. 3.85. Comparison of the oxidation kinetics for the El 10 unirradiated and irradiated claddings 


These data show that the oxygen weight gain in the irradiated cladding is somewhat higher than that in the 
unirradiated cladding, however, taking into account the sparsity of test data available for the El 10 irradiated 
cladding, it seems to be impossible to estimate this effect quantitavely. 

The El 10 oxidation kinetics was additionally analyzed to clarify the following important issues: 

• the assessment of the applicability of the El 10 oxidation kinetics (developed basing on the test data with 
fast heating and fast cooling modes) for the transient oxidation modes such as slow heating and slow 
cooling (S/S), slow heating and fast cooling (S/F), fast heating and slow cooling (F/S); 

• the comparison of the E635, Zry-4, and El 10 oxidation kinetics; 

• the comparison of the El 10 oxidation kinetics developed in different laboratories. 

The temperature-time histories of several tests performed at F/S, S/F, S/S combinations of heating and cool¬ 
ing rates were processed using the same procedure of the effective time determination, which was employed 
for this goal on the processing of F/F and F/Q appropriate data. The major provisions for the procedure are 
described in Appendix A-6 of the report. The comparative data characterizing the oxidation kinetics at dif¬ 
ferent oxidation modes are presented in Fig. 3.86. In accordance with the obtained data, reasonable correla¬ 
tions are observed between all types of experimental results. 


3.82 

















































































Fig. 3.86. The comparison of data characterizing the transient test modes 

with the E110 oxidation kinetics 


The oxidation kinetics sensitivity of zirconium niobium alloys to the alloying composition was verified using 
the data presented in Fig. 3.87. The comparison of the El 10 and E635 test data shows that the oxidation ki¬ 
netics of these two alloys are either the similar or very close to each other. 



Fig. 3.87. The comparative data on the El 10 and E635 oxidation kinetics 


The comparison of the El 10 and Zry-4 oxidation kinetics confirmed the results obtained by other researchers 
earlier. The El 10 oxidation rate is somewhat less than that for the Zry-4 cladding (see Fig. 3.88). 


3.83 
































































25 


20 - 


e 

u 

'Sb 

r-< 

c 


cS 
00 
-4—t 

-C 

oo 

"5 


15 


0 


1100 c 

Zry-4, Cathcart-Pawel [28] 
E110, RRC KI/RIAR 











k-zry4-l 100 -v2 


500 


1500 


1000 
Time (s) 

Fig. 3.88. The comparison of the El 10 and Zry-4 oxidation kinetics 


2000 


To estimate the agreement between the El 10 oxidation kinetics stated basing on results of this work with the 
results of other researchers, the appropriate data were compared on the basis of published investigation data 
obtained in the following organizations: 

• KFKI, Hungary [9]; 

• NFI, Czech republic [26]; 

• VNIINM, Russia [6]; 

• NC in Rossendorf, Germany [24], 

The comparison of the appropriate results performed at 1100 C allows to conclude that a good agreement is 
observed between the data of the RRC KI/RIAR (this work), NFI, VNIINM, and NC in Rossendorf (Fig. 
3.89). The KFKI data overestimate significantly the El 10 oxidation kinetics (may be due to end effects, see 
section 3.3.7). 



Fig. 3.89. The El 10 oxidation kinetics at 1100 C in accordance with the data 
obtained in different laboratories 


3.84 





















































References for Section 3 


[1] Atomic Energy Commission Rule-Making Hearing, Opinion of the Commission. Docket RM-50-1, 28 

December, 1973. 

[2] Bibilashvili Yu.K., Sokolov N.B., Andreeva-Andrievskaya L.N., Tonkov V.Yu., Salatov A.V., Morozov 

A.M., Smirnov V.P. "Thermomechanical Properties of Zirconium-Based Alloys Oxidized Claddings in 
LOCA Simulationg Conditions", Proc. of IAEA Technical Commettee Meet, on "Fuel Behavior under 
Transient and LOCA conditions ", Halden, Norway, September 10-14, 2001. 

[3] Baker L., Just L.C. "Studies of Metal Water Reactions at High Temperatures". Technical report of ANL- 

6548, 1962. 

[4] Solyany V.I., Bibilashvili Yu.K., Dranenko V.V., Izrailevskiy L.B., Levin A.Ya., Morozov A.A. "Char¬ 

acteristics of Corrosion Behavior of Zr-l%Nb WWER Fuel Claddings Within 700-1000°C on Long 
Term Exposures", Proc. of IAEA Technical Committee Meet, on "Water Reactor Fuel Behavior and Fis¬ 
sion Products Release in Off-Normal and Accident Conditions ", Vienna, 10-13 November, 1986. 

[5] Solyany V.I.. Bibilashvili Yu.K.. Dranenko V.V., Izrailevskiy L.B., Levin A.Ya.. Morozov A.A. "Studies 

of the corrosion behavior of Zr-l%Nb claddings in steam at high temperatures”, VANT. series " “ Mate¬ 
rials for Atomic Energy ", Vol.2(27), 1988, (in Russian). 

[6] Bibilashvili Yu.K.. Sokolov N.B.. Salatov A.V., Andreyeva-Andriyevskaya L.N., Nechayeva O.A., 
Vlasov F.Yu. "RAPTA-5 Code: Modelling of Behaviour of Fuel Elements of VVER Type in Design 
Accidents. Verification Calculations", Proc. of IAEA Technical Committee Meeting on "Behaviour of 
LWR Core Materials under Accident Conditions", Dimitrovgrad, Russia, on 9-13 October 1995. IAEA- 
TECDOC-921, Vienna, 1996. 

[7] Bibilashvili Yu.K., Sokolov N.B., Andreyeva-Andriyevskaya L.N., Salatov A.V. "High Temperature 
Interaction of Fuel Rod Cladding Material (Zr-l%Nb Alloy with Oxygen-Containing Mediums", Proc. 
of IAEA Technical Committee Meeting on "Behaviour of LWR Core Materials under Accident Condi¬ 
tions", Dimitrovgrad, Russia, on 9-13 October 1995. 

[8] Bohmert J. "Embrittlement of Zr-l%Nb at Room Temperature after High-Temperature Oxidation in 

Steam Atmosphere", Kemtechnik ,57 Nol, 1992 

[9] Gyori Cs. et.al. “Extension of Transuranus code applicability with niobium containing models 
(EXTRA)”. Proc. of FISA-2003 Conference EU Research in Reactor Safety’, EC Luxembourg, Novem¬ 
ber 2003. 

[10] Vrtilkova V., Valach M.. Molin L. "Oxiding and Hydrating Properties of Zr-l%Nb Cladding Material in 
Comparison with Zircaloys", Proc. of IAEA Technical Committee Meeting on "Influence of Water 
Chemistry on Fuel Cladding Behavior", Rez (Czech Republic), October 4-8, 1993. 

[11] Yegorova L„ Asmolov V., Abyshov G„ Malofeev V., Avvakumov A.. Kaplar E., Lioutov K.. Shestopa- 
lov A., Bortash A., Maiorov L., Mikitiouk K., Polvanov V., Smirnov V., Goryachev A., Prokhorov V., 
and Vurim A. “Data Base on the Behavior of High Bumup Fuel Rods with Zr-l%Nb Cladding and UO : 
Fuel (VVER Type) under Reactivity Accident Conditions”, RRC "Kurchatov Institute" report NSI RRC 
2179, Vol.1-3, 1999 (also USNRC report NUREG/IA-0156 and IPSN report IPSN 99/08 - 2). 

[12] Hobson D.O. and Rittenhouse P.L. "Embrittlement ofZircaloy Clad Fuel Rods by Steam During LOCA 
Transient", ORNL-4758, Oak Ridge National Laboratory, 1972. 

[13] Chung H.M. and Kassner T.F. "Embrittlement Criteria for Zircaloy Fuel Cladding Applicable to Acci¬ 
dent Situations in Light-Water-Reactors", NUREG/CR-1344, January 1980. 

[14] Uetsuka H. et.al. "Failure-Bearing Capability of Oxidized Zircaloy-4 Cladding under Simulated Loss- 
of-Coolant Conditions", J. Nucl. Sci. Tech. 20 (1983). 

[15] Bibilashvili Yu.K., Sokolov N.B., Dranenko V.V., Kulikova K.V., Izrailevskiy L.B., Levin A.Ya., 
Morozov A.M. "Influence of Accident Conditions due to Loss of Tightness by Primary Circuit on Fuel 
Claddines", Proc. of the Ninth Int. Svmp. on Zirconium in the Nuclear Industry , Kobe, Japan, November 
5-8, 1990. 


3.85 


[16] Bibilashvili Yu.K. et al.,"Influence of Accident Conditions due to Loss of Tightness by Primary Circuit 
on VVER-1000 Fuel Rod State", VANT, series “ Materials for Atomic Energy", Vol.2(42), 1991(rus). 

[17] Bbhmert J. "Embrittlement of Zr-l%Nb at Room Temperature after High-Temperature Oxidation in 
Steam Atmosphere", Journal, Kemtechnik 57 Nol, 1992. 

[18] Hozer Z., Griger A., Matius L., Vasaros L., Horvath M. "Effect of Hydrogen Content on the Embrittle¬ 
ment of ZR Alloys", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Transient 
and LOCA Conditions", Halden, Norway, September 10-14, 2001. 

[19] Kaplar E., Yegorova L., Lioutov K., Konobeyev A., Jouravkova N., Smirnov V., Goryachev A., Prok¬ 
horov V., Yeremin S., Svyatkin A. "Mechanical Properties of Unirradiated and Irradiated Zr-l%Nb 
Cladding", RRC "Kurchatov Institute" report NSI RRC 2241, 2001 (also USNRC report NUREG/IA- 
0199 and IPSN report IPSN 01-16). 

[20] Prokhorov V., Makarov O., Smirnov V., Goryachev A., Yegorova L., Kaplar E., and Lioutov K. 
“Improvement of method to measure tensile properties of Zr-l%Nb alloy cladding with simple ring 
specimens”. Proceedings of the 6' h inter-branch conference on reactor material science, Dimitrovgrad, 
Russia, September 11-15, Vol.2, pp. 209-212 (rus). 

[21] Bernstein M.L., Zaimovsky V.A. “Mechanical Properties of Metals”, Moscow, Metallurgy, 1979, (rus). 

[22] Zaimovsky A., Nikulina A., Reshetnikov N. “Zirconium alloys in the Nuclear Industry”, Moscow, En- 
ergoizdat, 1981 (rus). 

[23] Brachet J., Pelchat J., Hamon D., Maury R., Jaques P., Mardon J. "Mechanical Behavior at Room Tem¬ 
perature and Metallurgical Study of Low-Tin Zry-4 and M5™ (Zr-NbO) Alloys after Oxidation at 
1100°C and Quenching", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Tran¬ 
sient and LOC A Conditions", Halden, Norway, September 10-14, 2001. 

[24] Bohmert J., Dietrich M., Linek J. "Comparative Studies on High-Temperature Corrosion of Zr-l%Nb 
and Zircaloy-4", Nuclear Engineering and Design, 147 Nol, 1993. 

[25] Hozer Z. et. al. “Ring Compression Tests with Oxidised and Hydrided Zr-l%Nb and Zirkaloy -A Clad¬ 
ding”, Hungarian Academy of Sciences, CRIP, Budapest, Report KFKI-2002-01/G. 

[26] Vrtilkova V. et. al. “An Approach to the Alternative LOCA Embrittlement Criterion”, Proc. of SEGFSM 
Topical Meeting on LOCA Fuel Issues, Argonne National Laboratory, May 2004 
(NEA/CSNI/R(2004) 19). 

[27] Billone M.C., Yan Y. and Burtseva T. “Post-Quench Ductility of Advanced Alloy Cladding”, Nuclear 
Safety Research Conference (NSRC-2004), Washington, DC, October 25-27, 2004. 

[28] Cathcart J.V. and Pawel R.E. "Zirconium Metal-Water Oxidation Kinetics: IV. Reaction Rate Studies". 
ORNL/NUREG-17, 1977. 

[29] Chung H.M. “The Effects of Aliovalent Elements on Nodular Oxidation of Zr-based Alloys”, Proceed¬ 
ings of the 2003 Nuclear Safety Research Conference, Washington DC, October 20-22, 2003, 
NUREG/CP-0185, 2004. 

[30] Yegorova L., Lioutov K., Smirnov V., Goryachev A., Chesanov V. “LOCA Behavior of El 10 Alloy”, 
Proceedings of the 2003 Nuclear Safety Research Conference, Washington DC, October 20-22, 2003, 
NUREG/CP-0185, 2004. 

[31] Lemaignan C. “Physical Phenomena Concerning Corrosion Under Irradiation of Zr Alloys”, Zirconium 
in the Nuclear Industry: Thirteenth International Symposium, ASTM STP 1423, 2002, pp.20-29. 

[32] Uetsuka H. et al. "Zircaloy-4 Cladding Embrittlement due to Inner Surface Oxidation under Simulated 
Loss-of-Coolant Conditions", J. Nucl. Sci. Tech. 18 (1981). 

[33] Uetsuka H. et al. "Embrittlement of Zircaloy-4 due to Oxidation in Environment of Stagnant Steam", J. 
Nucl. Sci. Tech. 19(1982). 

[34] Mardon J.P. et al. “Update on the Development of Advanced Zirconium Alloys for PWR Fuel Rod 
Claddings”, Proceedings of the International Topical Meeting on Light Water Reactor Fuel Perform¬ 
ance, Portland, Oregon, March 2-6, 1997. 


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[35] JC. Brachet, L. Portier, V. Maillot, T. Forgeron, JP. Mardon, P. Jacques, A. Lesbros, “Overview of the 
CEA data on the influence of hydrogen on the metallurgical and thermal-mechanical behavior of Zir- 
caloy-4 and M5™alloys under LOCA conditions”, Nuclear Safety Research Conference 

(NSRC-2004), Washington. DC, October 25-27, 2004. 

[36] Mardon J., Frichet A., Bourhis A. "Behavior of M5™ Alloy under Normal and Accident Conditions", 
Proc. of Top Fuel-2001 Meet., Stockholm. May 27-30, 2001. 

[37] Asmolov V., Yegorova L.. Lioutov K., Smirnov V., Goryachev A., Chesanov V.. Prokhorov V. "Under¬ 
standing Loca-Related Ductility in El 10 Cladding", Proceedings of the 2002 Nuclear Safety Research 
Conference, Held at Marriott Hotel at Metro Center Washington, DC, October 28-30, 2002, 
NUREG/CP-0180, pp. 109-125. 

[38] Yegorova L., Lioutov K.. Smirnov V. “Major Findings of El 10 Studies under LOCA Conditions”, Proc. 
of SEGFSM Topical Meeting on LOCA Fuel Issues, Argonne National Laboratory, May 2004 
(NEA/CSNI/R(2004) 19). 

[39] Pirogov Ye.N., Alimov M.I.. Artiuchina L.L. “Creep of H-l alloy in the range of phase transition” Jour¬ 
nal of Soviet Atomic Energy, v.65 (3), p. 293, 1988 (rus). 

[40] Forgeron T. et. al, “Experiment and Modeling of Advanced Fuel Rod Cladding Behavior Under LOCA 
Conditions: Alpha-Beta Phase Transformation Kinetics and EDGAR Methodology”, Zirconium in the 
Nuclear Industry: Twelfth International Symposium,. ASTM STP 1354, 2000, pp. 256-278. 

[41] Nikulina A. et.al. “Zirconium Alloy E635 as a Material for Fuel Rod Cladding and Other Components 
of WER and RBMK Cores”, Zirconium in the Nuclear Industry': Eleventh International Symposium , 
ASTM STP-1295, November 1996. 

[42] Douglass D.L. “The Metallurgy of Zirconium”, International Atomic Energy Agency, 1971. 

[43] Garde A.M. “Influence of Cladding Microstructure on Low Enthalpy Failures in RIA Simulation Tests”, 
Zirconium in the Nuclear Industry': Twelfth International Symposium, ASTM STP 1354, 2000. 

[44] Hache G., Chung H. "The History of LOCA Embrittlement Criteria", Proc. of the 28 h Water Reactor 
Safety Information Meeting, NUREG/CP-0172, 2000. 

[45] Bai J., Prioul C., Pelchat J., Barcelo F. “Effects of Hydrides on the Ductile Brittle Transition in Stress- 
Relieved, Recrystallized and ^-Treated Zircaloy”, Proc. of the International Topical Meeting on the 
LWR FuelPerfonnance, Avignon. France, 1991. 

[46] Waeckel N., Mardon J.P. “Recent data on M5™Alloy under LOCA Conditions”. Proceedings of the 
2003 Nuclear Safety Research Conference , Washington DC, October 20-22, 2003, NUREG CP-0185, 
2004. 

[47] G. Feamehough and A. Cowan “The effect of hydrogen and strain rate on the “ductile-brittle” behavior 
of Zircaloy”, Journal of Nuclear Materials NE22, 1967. 

[48] F. Nagase, K. Ishijima. and T. Furuta, “Influence of Locally Concentrated Hydrides on Ductility of Zir- 
caloy-4”, Proc. of the CSNI Specialist Meeting on Transient behaviour of High Burnup Fuel, Cadarche, 
France, September 12-14, 1995. 


3.87 









































4. Oxidation behavior and embrittlement threshold 

OF THE MODIFIED E110 CLADDING: PROGRAM AND DISCUSSION 
OF TEST RESULTS 

4.1. Major provisions of the test program with a modified El 10 cladding 

The results of the test program performed with the El 10 standard as-received tubes and presented in chapter 
1 of the report have shown that: 

• the zero ductility threshold of the El 10 alloy is lower than that of the Zry-4 alloy; 

• the earlier initiation of the breakaway oxidation accompanied by hydrogen absorption in the prior 
(3-phase of the El 10 cladding is the major reason for the different mechanical behavior of oxidized 
claddings fabricated of these two alloys. 

The numerous sensitivity studies performed in the frame of this work allowed to establish that variations in 
oxidation modes may somewhat hasten or delay the initiation of the breakaway oxidation in the El 10 alloy, 
however, any variations in test conditions do not allow to avoid completely this type of oxidation in the ECR 
range of interest. 

Approximately the same conclusion was made basing on the analysis of results of tests performed to check 
the correlation between some alloying elements (O, Fe, Sn) and the oxidation behavior of niobium- 
containing claddings. 

All these results allowed to report the following general question: is this oxidation behavior typical of the 
whole family of niobium-bearing alloys or only of the El 10 alloy? The comparison of the El 10 test results 
with the published data on the other Zr-l%Nb alloy, namely, the M5 alloy [1, 2], allowed to reveal the 
following: 

• the M5 and Zry-4 claddings are characterized by the similar embrittlement threshold at 1100 C; 

• the breakaway oxidation was not observed on the M5 claddings in the tested ECR range. 

A careful analysis of the existing situation allowed to assume that the specific behavior of the El 10 cladding 
may be the consequence of tw o groups of effects connected with the cladding fabrication: 

1. Surface effects. 

2. Bulk effects. 

The surface effects combine such factors as the surface chemistry (surface contaminations) and surface 
roughness. In their turn, bulk effects may be the manifestation of such factors as the cladding material 
chemical composition (the composition of impurities) and microstructure effects (the phase composition, 
grain size, parameters of the secondary precipitates, etc.). 

As for the surface effects, it is known that the surface finishing of cladding tubes is the important component 
for the cladding corrosion resistance. Thus, for illustration, the results of this work confirm the numerous 
observations concerning the relationship between the initiation of the breakaway oxidation and the localiza¬ 
tion of the cladding surface scratches (see Fig. 4.1). 

To provide the chemical cleaning and chemical polishing of the El 10 cladding tube, the standard procedure 
for the E110 surface finishing is used. This procedure includes the chemical etching and anodizing of the 
outer surface of the El 10 fuel rod. Besides, special studies are performed at present to develop the modified 
method for the El 10 surface finishing. One of the potential variants of new approaches to this problem is the 
grinding of the cladding outer surface and the jet etching of the cladding inner surface. Taking into account 
all these considerations, it was decided to perform several special tests the assess the sensitivity of the El 10 
oxidation behavior to surface effects. 


4.1 



Sample #27 

ECR=4.6% 


surface scratches on as-received El 10 cladding sample 


Fig. 4.1. The relationship between surface scratches and localization of the breakaway oxidation areas 


The studies of bulk effects associated with the oxidation behavior of niobium-bearing alloys were the subject 

of numerous investigations performed in Russia and other countries. As for the microstructure effects, the 

analysis of publications allowed to note the following: 

• the corrosion properties of binary zirconium-niobium alloys depend directly on the content and phases of 
the zirconium and niobium [3]; 

• the formation of equilibrium a-Zr and P-Nb phases leads to the improvement of corrosion resistance of 
binary alloys [3]; 

• the corrosion properties of the El 10 alloy depend on the material microstructure; the annealing tempera¬ 
ture and duration are the major factors determining the cladding microstructure; the annealing at 580 C 
during 3 hours (below the monoeutectoid line) leads to the maximum depletion of the a-Zr matrix by the 
niobium and to the formation of the P-Nb phase that improves the alloy corrosion resistance [4]; 

• the corrosion behavior of the El 10 alloy cladding is shown to be strongly dependent on the material 
structure (including the grain size), the formation of the completely recrystallized material with a fine 
grain size determines the best corrosion resistance [5]; 

• a strong dependence of corrosion properties of binary alloys on heat treatment was revealed; the forma¬ 
tion of metastable phases at the annealing temperature higher than the monoeutectoid one leads to the 
very low corrosion resistance, the annealing in the temperature range of the a-Zr phase presence leads to 
the formation of the equilibrium structure and high corrosion resistance [6]; 

• to prevent the degradation in the corrosion resistance, the intermediate and final annealing temperature 
must not exceed 600 C; in this case, the material of M5 tubes remains in the a-Zr plus P-Nb region under 
completely recrystallized conditions [7]; 

• the samples of Zr-xNb binary alloys annealed at 570 C to form P-Nb phase showed a much lower corro¬ 
sion rate and higher volume fraction of tetragonal Zr0 2 in the oxide than those annealed at 640 C to form 
p-Zr phase [8]; 

• basing on the consideration of publications devoted to the corrosion behavior of niobium-bearing alloys, 
the authors of investigations presented in [9] have declared that the correlation between the oxide charac¬ 
teristics and microstructure or microchemistry, such as P phases, Nb-containing precipitates and soluble 
Nb in the matrix, is not well understood in Zr-Nb alloys. One of the major results of investigations per¬ 
formed by these researchers could be formulated as following: at high Nb-contents 1.0-5.0 wt%, the cor¬ 
rosion rate was very sensitive to the annealing condition, the transformation of the oxide structure from 
tetragonal Zr0 2 to monoclinic Zr0 2 and of the oxide microstructure from the columnar to equiaxed struc¬ 
ture was accelerated in the samples having p-Zr phase. The corrosion rate of samples annealed at 570 C 
with the formation of the P-Nb phase was much lower. Moreover, it is assumed that the equilibrium con¬ 
centration of Nb in the a-matrix would be a more dominant factor to enhance the corrosion resistance 
that the Nb-containing precipitates. 

The above listed results of previous investigations have shown that the corrosion behavior of binary Zr-Nb 

alloys is very sensitive to the microstructure characteristics that in their turn are determined by the 

4.2 













fabrication procedures and conditions. In this case, the fact cannot be excluded that the annealing condition 
may be not the only key factor in spite of the fact that recent special studies performed with the M5 alloy 
have demonstrated that the corrosion resistance is not sensitive to many aspects of the fabrication process [7]. 
L ntortunately, a direct comparison of manufacturing procedures used on the fabrication of the El 10 cladding 
and M5 cladding to reveal the differences that may be responsible for differences in the corrosion behavior 
of these two niobium-bearing claddings is not possible because this is confidential commercial information. 
Therefore, to provide the comparative analysis of possible microstructure effects in these two alloys, it was 
decided to perform a careful SEM (scanning electron microscopy) and TEM (transmission electron 
microscopy) examinations of the El 10 cladding material and after that to compare the obtained results with 
the published data characterizing the M5 tube microstructure. 

The next position of special investigations performed with the El 10 cladding to study of bulk effects was 
devoted to the studies of the microchemical effects of the cladding material in the context of the El 10 oxida¬ 
tion behavior. The analysis of results of previous investigations performed in this line allowed to reveal the 
following: 

• the optimization of Nb concentration in the El 10 alloy allowed to minimize the influence of N, Al, C 
impurities causing the degradation of the corrosion resistance [4]: 

• the secondary phase particles influence significantly the corrosion behavior of niobium-bearing alloys. 
Special investigations performed with the cladding samples the surface of which was modified due to the 
ion alloying have shown that a strong dependence is between the oxide structure and secondary intermet- 
allic participates on the cladding surface. So. the oxide layer has a thickening in this area consisting of 
metal oxides of the secondary phase participates. The decrease of the participate density at the alloying 
process leads to the improvement of the oxide contact with the metal matrix [10]; 

• the corrosion tests have demonstrated that the alloying of the El 10 cladding of such elements as Fe, Mo, 
Sn leads to the decrease in the oxidation rate [11]; 

• the presence of such elements as Sn. Fe. Cr, Nb improves the corrosion resistance of zirconium [6]. 

It is evident that the above listed results of previous investigations are remarkable for the fragmentariness 
and are insufficient to analyze microchemical bulk effects as applied to the corrosion behavior of the El 10 
alloys. The additional analysis of published data devoted to the relationship between the corrosion behavior 
of other zirconium alloys and their microchemical compositions has shown that in spite of a great number of 
investigations performed recently, a clear physical understanding of appropriate phenomena has not been 
achieved yet. Taking into account this circumstance, the further development of this line of investigations 
was based on the following considerations; 

• if the microchemistry of Zr-Nb alloys is the key factor resulting in differences in the behavior of such 
two Zr-l%Nb alloys as El 10 and M5. then this means that these two alloys have different compositions 
of impurities in the cladding material. In this case, this conclusion does not concern the differences in the 
oxygen concentration as this effect has been specially investigated (see chapter 1); 

• the analysis of possible reasons for potential differences in the impurity composition and content of these 
two alloys has shown that two different methods are employed to fabricate the Zr-l%Nb ingot: 

- the sponge Zr is used on the manufacturing of the M5 cladding; 

- the mixture of iodide and electrolytic Zr is used on the manufacturing of the E110 cladding; 

• the comparative description of these methods for the preparation of the Zr-l%Nb ingot is given in [12]. 

It should be noted that the Russian industry is planning in the nearest time to proceed to the employment of 
sponge Zr to manufacture the El 10 cladding taking into account the economic advantage (the fabrication of 
iodide/electrolytic Zr is significantly more expensive than the fabrication of sponge Zr). In the frame of this 
work, different types of the El 10 claddings were manufactured with the use of sponge Zr. Therefore, it w r as 
decided to perform the investigations of microchemistry effects associated with the oxidation behavior of the 
E110 alloy using several different types of the El 10 cladding manufactured on the basis of sponge Zr. 

In accordance with the above listed surface and bulk effects, the program of this stage of research w as devel¬ 
oped. The major provisions of this program are presented in Table 4.1. 


4.3 


Table 4.1. The El 10 surface and bulk effects studies: major provisions of experimental program 





1. Surface effects 
investigations 

1.1. Tests of the El 10 
etched and anodized 
cladding 

To perform several oxidation tests with the 
E110 etched and anodizing cladding and to 
compare the mechanical behavior of the 

E110 cladding and El 10 as-received tube 

The commercial El 10 etched 
and anodizing cladding should 
be used for these tests 

1.2. Tests of the El 10 
cladding with the 
modified surface 
finishing 

To perform oxidation tests with two types of 
surface finishing: 

1. Grinding of the cladding outer surface and 
jet etching of the cladding inner surface 

2. Polishing of the cladding outer and inner 
surfaces 

The commercial advanced El 10 
cladding should be used for the 
first type of tests. The 
laboratory procedure for the 
preparation of special samples 
with polished outer and inner 
surfaces of the El 10 as-received 
tubes should be used for the 
second type of tests 

2. Bulk effect 
investigations: 

2.1. Studies of the E110 
cladding microstructure 

To perform the SEM and TEM examinations 
of different types of the El 10 claddings and 
to perform the comparative analysis of 
obtained microstructure characteristics 

The samples of the El 10 
cladding material fabricated 
with the application of different 
methods for manufacturing of 
Zr-l%Nb ingot should be 
prepared for these examinations 

2.2. Studies of 
dependence of the 
oxidation and mechanical 
behavior of the El 10 
cladding on the chemical 
composition of Zr-l%Nb 

To perform oxidation tests with several types 
of the El 10 and E635 as-received tubes 
fabricated with the use of sponge Zr-1 %Nb. 

To compare the mechanical behavior of these 
claddings with the behavior of standard El 10 
tubes manufactured with the use of 
iodide/electrolytic Zr 

Special cladding samples 
fabricated in accordance with 
the standard procedure but with 
the application of sponge Zr 
should be used for these tests 


All oxidation and mechanical tests performed in the frame of this program were made in accordance with the 
following technical requirements: 

• the oxidation type: the double-sided oxidation with the F/F combination of heating and cooling rates; 

• the oxidation temperature: 900-1200 C; 

• the characterization of mechanical tests: ring compression tests of 8 mm rings at 20 and 135 C. 

4.2. The analysis of experimental results obtained at surface effect studies 

As it was noted in section 4.1, three types of cladding samples were used to investigate these effects: 

1. Standard E110 etched and anodized cladding (E110A). 

2. E110 as-received tube with the modified surface finishing (the outer surface grinding and inner surface 
jet etching) performed at the tube plant (El 10 m ). 

3. E110 as-received tube sample, one half of which remained in the initial state and the second half was 
polished from the inside and outside in the RIAR laboratory (El 10 po |). 

The whole scope of obtained test results is presented in Tables B-l and B-2 of Appendix B. The appearances 
of the El 10A and Ell 0 m samples before and after oxidation tests are shown in Fig. 4.2. 


4.4 














before oxidation 


View A 


E110A (#117) 
ECR-9.0% 

H 2 content=550 ppm 



after oxidation 


!a 



before oxidation 


View A 



Ell 0 m (#136) 
ECR=7.7% 

H 2 content=90 ppm 





Fig. 4.2. The appearance of the El 10 claddings fabricated with the use of two different types 
of surface finishing before and after oxidation tests at 1100 C 


The analysis of these data allows to note the following: 

• the standard surface finishing of the El 10 cladding does not lead to improvements in comparison with 
the oxidation behavior of the El 10 as-received tubes; the classic breakaway effect accompanied by the 
hydrogen pickup was observed on the cladding outer surface (see Fig. 4.3 also); 


• the modified surface finishing of the El 10 tube on the basis of the outer surface grinding and inner sur¬ 
face jet etching does not allow to eliminate the breakaway oxidation. The oxidized cladding appearance 
and the microstructure of oxides formed on the inner and outer cladding surfaces (see Fig. 4.3) show that 
the spallation and delamination of Zr0 2 oxide are observed on both surfaces of the cladding. Moreover, 
the oxide thickness on the etched inner surface is larger than that on the grinded outer surface. 


El 10A (etched and anodized cladding outer surface) 
Outer surface 




Ell 0 m (grinded outer surface and jet etched inner surface) 



Fig. 4.3. The microstructure of the El 10A and El 10 m claddings after the oxidation at 1100 C 


4.5 


































Obtained observations are in the reasonable agreement with results of ring compression tests performed with 
the 8 mm samples cut off from 100 mm oxidized samples (see Fig. 4.4). Moreover, numerous special inves¬ 
tigations performed in the ANL with the El 10 samples which underwent different procedures of the cladding 
etching and the cladding cleaning after etching have demonstrated that etching is more than undesirable pro¬ 
cedure from the point of view of the El 10 cladding oxidation behavior [13]. This conclusion is confirmed by 
the results of previous studies of zirconium alloys [14]. Taking into account a very high sensitivity of the 
breakaway initiation to the minimum presence of F-contamination on the cladding surface (that was demon¬ 
strated many times in previous investigations), it may be assumed that fluorine containing particles remain 
on the etched cladding surface, though this assumption contradicts the results of the appropriated check pro¬ 
cedures that are performed at tube plant. Unfortunately, it was impossible to perform special measurements 
of fluorine contents on the cladding surface in the frame of this work as those are very difficult from the 
methodological point of view. 


V? 


O 

3 

”c3 

3 

3 

oz 


70 

60 

50 

40 

30 


10 

0 







V 

\ 

\ 

\ 






A El 10m 

• E110A 

- as-received 



\ 

\ 

\ 

\ 

\ 




\ 

\ 

X 

\ 

X 








V 

X 

X 

\ 

\ 




X 

X 

X 

X 

X 




A 

- 4 

\ 

\ 

\ 

-- 

el lOsh-en 


4 6 8 10 12 

ECR (%) 

Fig. 4.4. The comparison of residual ductility of the standard El 10 as-received tube, E110A and E110 r 

claddings after the oxidation at 1100 C 


The indirect confirmation of this reason for the El 10 breakaway initiation was obtained from the results of 
investigations performed in the ANL. They performed additional polishing of the outer and inner surfaces of 
the El 10A cladding. In this case, the cladding inner diameter was increased after polishing by 100 pm. The 
oxidation tests at 1100 C and ring compression tests have shown that the margin of residual ductility remains 
in these claddings up to 16% ECR (as-measured) [13]. 

The similar tests were performed in the frame of this work also. Polishing of one half of the El 10 sample 
100 mm long was performed on the inner and outer surfaces of the El 10 as-received tube. The second half of 
this sample remained intact. The surface roughness on the polished part of the cladding sample was about 
0.08-0.16 pm that corresponds to the surface roughness of the M5 and Zry-4 claddings (0.1-0.15 pm). Tak¬ 
ing into account the effect of the oxidation temperature revealed during the first part of research program, the 
oxidation tests of these samples were carried out at 1000 and 1100 C. The organized tests are presented in 
Fig. 4.5. The tabular results of tests are listed in Tables B-l and B-2 of Appendix B. 


4.6 





























Fig. 4.5. Demonstration of sensitivity of the oxidation and mechanical behavior for the El 10 cladding 

to the cladding surface polishing 


The analysis of obtained comparative data allowed to reveal the following important aspects of the El 10 
oxidation behavior as a function of the cladding surface machining: 

• the rate of the El 10 cladding unpolished part oxidation is higher than that of the polished region; 

• indications of the breakaway oxidation including the hydrogen absorption practically disappear on the 
polished part of the El 10 cladding. This process takes place especially clearly on the oxidation at 
1000 C; 

• the residual ductility of the El 10 cladding polished part is many times increased in comparison with the 
oxidized cladding unpolished part. 

The additional information to characterize the appropriate cladding behavior may be obtained due to the con¬ 
sideration of the oxidized cladding microstructure in the polished and unpolished parts (see Fig. 4.6). As can 
be observed, the oxide spallation is noted on the outer surface of cladding unpolished part. The contact of 
oxide with the cladding matrix on the inner side is somewhat better but it is seen that a systematical layer of 
pores has been already generated on the interface between the oxide and metallic matrix, and. consequently, 
the oxide layer spallation will take place at the ECR increase. Another case is observed on the oxidized clad¬ 
ding polished part. The uniform oxide layer having a good adhesion with the metallic matrix covers the clad¬ 
ding outer and inner surfaces. This oxide behavior provides the lower oxidation rate and low rate of hydro¬ 
gen diffusion into the El 10 cladding. 


4.7 





















































Fig. 4.6. The oxide microstructure on the polished and unpolished parts of the El 10 cladding 

after the oxidation at 1000 C 

The same results were obtained from the tests with polished El 10 claddings in the ANL [13]. Polishing of 
the E110 cladding surface allowed to delay greatly the breakaway initiation, that in its turn resulted in the 
significant improvement of the oxidized cladding mechanical properties. 

The results of tests with polished El 10 claddings do not allow to formulate the final conclusion concerning 
the mechanism due to which the El 10 cladding oxidation behavior is improved. However, turning back to 
the consideration of two possible reasons for the surface effect (the surface roughness and surface 
contamination) and taking into account the obtained test data, the following may be assumed: 

• both factors are important for the initiation of the breakaway oxidation; 

• but even an ideal surface state from the point of view of the surface roughness does not allow to avoid 
the breakaway oxidation at the low ECRs in the presence of the surface contaminations that is confirmed 
by the tests with etched claddings. 

Thus, the microchemistry of the cladding surface is apparently the dominant factor influencing the break¬ 
away condition. The further analysis of microchemical effects will be continued in the next section of the 
report. 


4.8 






























4.3. The assessment of relationship between the microchemical composition and 
oxidation behavior of niobium-bearing alloys 

To study these potential etfects ot the cladding behavior, several types of zirconium niobium claddings were 
selected tor the oxidation and mechanical tests. The specification for the used cladding material is presented 
in Table 4.2. The additional information characterizing the cladding initial parameters is given in Tables A-l, 
A-2 of Appendix A. 


Table 4.2. The specification for the used cladding material 


Notation conventions of cladding 
types used in this program 

Alloying composition 

Input components used on the 
fabrication of the alloy ingot 

E110 

Zr-1 %Nb 

Iodide Zr, electrolytic Zr, recycled 
scrap, Nb 

E 1 1 Oo(fr) 

Zr-1 %Nb 

French sponge Zr (CEZUS), Nb 

E 1 10 G (3fr) 

Zr-1 %Nb 

French sponge Zr, iodide Zr, recycled 
scrap, Nb 

E 1 1 0o(3ru) 

Zr-l%Nb 

Russian sponge Zr, iodide Zr, recycled 
scrap, Nb 

E 1 1 Oiow Hf 

Zr-1 %Nb 

Iodide Zr, electrolytic Zr with low Hf 
content, recycled scrap. Nb 

E635 

Zr-1 %Nb-1,2%Sn-0.35%Fe 

Iodide Zr, electrolytic Zr, recycled 
scrap, Nb, Sn, Fe 

E635o(ff) 

Zr-1 %Nb-1.2%Sn-0.35%Fe 

French sponge Zr (CEZUS), Nb, Sn, Fe 


All types of these claddings were manufactured in accordance with the Russian process for the cladding fab¬ 
rication. The subprogram of experimental investigations with the cladding material manufactured with the 
use of sponge Zr and electrolytic Zr with low Hf contents consisted of positions listed in Table 4.3. 


Table 4.3. The subprogram major tasks for the test with the sponge cladding material and El 10 cladding 

with low Hf content 


Task 

Motivation 

1. To perform the oxidation tests at 1100 C of the El 10 and 
E635 claddings manufactured with the use of 100% sponge 

Zr (El 10 G (ff), E635 G (f T )) and to perform the ring compression 
tests of oxidized claddings 

To develop the comparative test data on the 
behavior of iodide/electrolytic El 10 and 

E635 alloys and sponge El 10 and E635 
alloys 

2. To perform the tests with the El 10 cladding manufactured 
by the standard Russian procedure but with low Ht content in 
the electrolytic Zr (El 10| OW Hf) 

Taking into account that the chemical 
composition of the standard El 10 cladding is 
characterized by a very high content of Hf 
(in comparison with the sponge material), to 
check the dependence of the El 10 oxidation 
behavior on Hf content 

3. To perform the oxidation tests at 1100 C of the El 10 
claddings manufactured from the mixture of iodide, sponge 

Zr and recycled scrap, after that to estimate the mechanical 
behavior of oxidized samples (El lOoono.pfr)) 

To clarify the sensitivity of the cladding 
oxidation behavior to the variation in the 
microchemical composition of the El 10 
alloy 


4.9 


















Task 

Motivation 

4. To perform the oxidation tests at 900, 1000, 1200 C with 
the use of sponge types of the El 10 cladding material and to 
estimate the margin of residual ductility of oxidized samples 

To determine the representativity of test 
results obtained at 1100 C for other 
temperature regions 


The whole scope of test results obtained in the frame of this subprogram is presented in Tables B-l, B-2 of 
Appendix B and in Appendix H of the report. The major outcomes of the subprogram are discussed in the 
next paragraphs of this section. 


4.3.1. The analysis of the oxidation and mechanical behavior 
for the E110 G (fr) and E635 G(fr) claddings fabricated 
on the basis of 100% French sponge Zr 

Two samples from this type of the El 10 cladding were oxidized at 1100 C. The appearances of samples after 
the oxidation tests are presented in Fig. 4.7. 


Sample #89: 
10.5% ECR 





22 ppm of H content 


Sample #90 
13% ECR 



30 ppm of H content 


Fig. 4.7. Appearances of E110 G <f r ) samples after the double-sided oxidation at 1100 C 


The appearance of these samples demonstrates that in spite of a high level of measured ECRs (10.5 and 
11.0%) the cladding surface is covered with the black bright oxide. This fact indicates that the mechanism of 
the uniform oxidation was realized on testing of these samples. The results in detail of mechanical tests with 
this type of the El 10 cladding in comparison with the results obtained in the test with the standard El 10 
cladding allow to note the following (see Fig. 4.8): 

• the El 10 G (fr) oxidized rings have a visible residual plastic deformation after the ring compression tests; 

• the load-displacement diagrams of the El 10c(f r > oxidized samples demonstrate that a significant part of 
the sample deformation before the fracture was accompanied by the plastic strain; 

• both El 10 G (fr) oxidized samples are characterized by a very low hydrogen concentration; 

• the Ell 0 G(fr) oxidized claddings have a significant margin of residual ductility at 13% ECR (as- 
measured); this margin corresponds to that for the Zry-4 cladding. 

This first stage of appropriate investigations has shown that El 10 claddings fabricated in accordance with the 
traditional Russian method of ingot preparation and El 10 claddings fabricated in accordance with the west¬ 
ern method of ingot preparation have quite a different oxidation behavior. 

In the context of investigations presented in this section of the report, the revealed differences may come 
from the differences in the microchemical composition (impurity content) of the El 10 and El 10 G(fr) cladding 
materials. To extend the data base for the analysis of these effects, intermediate compositions of cladding 
materials such as El 10 G(3f t) and El lOc^n,) were tested. 


4.10 



















Appearance of ring sam- 
ples after mechanical tests 

#89 




Fig. 4.8. Results of ring compression tests with E110 G (f r ) oxidized samples 


4.3.2. The interpretation of test results with EllOcpfr) and EllOcpru) claddings 

The appearances of these samples after the oxidation test in comparison with the standard El 10 cladding and 
the results of mechanical tests are shown in Fig. 4.9 for an oxidation temperature of 1100 C. 

The obtained data showed that all effects revealed in the analysis of the EIIOg^) behavior (a black bright 
oxide, low hydrogen content, high margin of residual ductility up to 16.7% ECR), were confirmed in the 
tests of the EllOoptr) and EllO^ru) claddings. This observation indicates that in spite of the fact that 
El 10(3936-) and El lOooru) claddings contain only 70% of sponge Zr-l%Nb this turned out to be enough to 
improve considerably the oxidation behavior of the El 10 cladding. 

The additional confirmation of this conclusion can be obtained due to the comparative analysis of the clad¬ 
ding microstructures presented in Fig. 4.10. As can be seen from this data, the prior (3-phase of the cladding 
samples manufactured from the iodide/electrolytic alloy and oxidized to the 10.0% ECR (#65) was so em¬ 
brittled that local areas of the prior (3-phase flaked off on polishing of the metallographic sample (black areas 
on the sample surface). Besides, this sample has no clearly marked boundary between the a-Zr(O) and prior 
(3-phases. In contrast to this cladding, the cladding manufactured on the basis of sponge Zr (#89) is character¬ 
ized by a clear boundary line between a-Zr(O) and prior p-phases and the structure of the prior P-phase 
without visible indications of a-Zr(O) phase and solid hydrides. 

As a whole, these differences indicate that the diffusion processes in these two types of the El 10 cladding 
determining the diffusion of alloying elements (including such minor alloying elements as impurities), pro¬ 
ceeded in these cladding in different ways. It may be assumed that these differences represent the key factor 
predetermining the differences in the oxygen and hydrogen pick up and the distribution in the El 10 standard 
and E110 sponge claddings. At this stage of the preliminary analysis of revealed effects the so-called “haf¬ 
nium issue” arose. 

The basis for the hafnium issue lies in the fact that hafnium concentration in the El 10 standard alloy (up to 
500 ppm) and hafnium concentration in the El 10 alloy fabricated with the use of sponge (<100 ppm) differ 
greatly. Therefore, in spite of the fact that the difference in the Hf concentration cannot cause the general 
difference in the oxidation behavior of these cladding from the physical point of view, it was decided to per¬ 
form special investigations to study this effect. 


4.11 



























Fig. 4.9. Comparative data base characterizing the E110 G (3f r ) and E 110 G( 3 ru) 

oxidation/mechanical behavior 



100 mn 




§ 

'*vj. 1 


Iodide/electrolvtic El 10 


Sponge E110 


#68-5 

ECR=10.0% 
(El 10) 


Etched 


Etched 


#89-4 

ECR=10.5% 
(El 10 G(fi .)) 


Fig. 4.10. The comparison of microstructures for iodide/electrolytic and sponge El 10 claddings 

after the oxidation at 1100 C 

Two cladding samples manufactured with the employment of electrolytic Zr with low Hf content (90 ppm) 
were used for this goal. The results of tests performed with these samples are presented in Fig. 4.11, in Ta¬ 
bles B-l, B-2 of Appendix B and Fig. H-15 of Appendix H. 


4.12 



















































o 

rj 


a 

3 



O 

S' 


i 



ECR (%) 


Fig. 4.11. The appearance of the El 10| OW | lf cladding after the oxidation at 1100 C 

and results of mechanical tests 


The analysis of obtained data allows to state the following: 

• the decrease of Hf content in the Zr-l%Nb alloy improves the oxidation behavior of the El 10 cladding; 
so, the first indications of the breakaway oxidation were fixed at 9.2% ECR; 

• the Ell0| O w Hf sample oxidized up to 9.2% ECR has a significant margin of residual ductility and low 
hydrogen content (17 ppm); 

• the zero ductility threshold of the Ell 0 iow H f cladding is increased up to 12% ECR but the breakaway 
oxidation was observed in the ECR range 9-12% and the embrittlement of the El 10 cladding was caused 
by oxygen and hydrogen pickup (the hydrogen concentration was about 430 ppm) at 11-12% ECR. 

Therefore, it can be assumed that the process of electrolytic Zr cleaning of hafnium was associated with a 
change in the level of other chemical impurities. In this case, this change led to some improvement of the 
E110 oxidation behavior but not so radically as it took place for sponge types of the El 10 cladding. 
Therefore, further investigations were continued with the Ell0o(fr) and EllOcom) claddings to adjust the 
sensitivity of the oxidation behavior of these claddings to the oxidation temperature. 

4.3.3. The sensitivity of the oxidation behavior of the sponge El 10 cladding to 
the oxidation temperature 

The temperature range 900-1200 C was studied in these investigations. Taking into account the results of 
tests with the traditional El 10 cladding that showed that the oxidation behavior of the El 10 alloy at 1000 C 


4.13 








































was somewhat worse than at 1100 C, the first stage of temperature dependent tests was performed at 1000 C. 
These tests allowed to reveal the unexpected effect associated with a sharp decrease in the oxidation rate of 
the sponge type cladding at 1000 C. The scale of this effect can be characterized using the following 
experimental data: 

• it takes 865 seconds to oxidize the standard El 10 cladding up to 7.7% ECR; 

• it takes 2519 seconds to oxidize the sponge type of the El 10 cladding up to 6.9%; 

• and it takes 5028 seconds to oxidize the sponge type of the El 10 cladding up to 8.9% ECR. 

The oxidation kinetics of different types for the El 10 cladding will be considered in detail in the next 
paragraphs of the report. 

As for this paragraph, the following notice concerning the oxidation rate problem should be made: a very low 
oxidation rate of sponge El 10 cladding at 1000 C led to the fact that the time limit for the steam generator 
used in this test series (approximately 5000 s) was exhausted at the ECR of about 8.5-8.9%. Thus, this ECR 
range was the maximum one in the oxidation tests at 1000 C. The results of tests at 1000 C are presented in 
Fig. 4.12. 


4.14 


E110 

(tradi¬ 

tional) 


E110 


G(fr) 


ECR=7.7% 


ECR=6.5% 
(28 ppm of H 
content) 


t e f=865 s 


t e r=201 6 s 


#44 



#91 


E110 


G(3ru) 


ECR-6.9% 
(16 ppm of H 
content) 


t et =2519 s 



#98 


E110 


G(3ru) 


ECR=8.9% 
(11 ppm of H 
content) 


t e t—5028 s 



#101 


E110 


G(ir) 


ECR=8.5% 
(12 ppm of H 
content) 


t e f=5013 s 



#93 


i 


80 


g 60 

o 

3 40 


a 

a 


OL - 


20 


• o 


□ * 
i 

□ \ 


-E110, 1100C 

O El 10, 1000 C 
D EllO^fr,, 1000 C 
A El lO^j^, 1000 C 


nmnc-\:?rf 


—r- 

4 


—I— 

6 


±LL 


8 10 
ECR (%) 


12 


14 


16 


Fig. 4.12. The appearance and mechanical properties of different El 10 claddings 

after the oxidation at 1000 C 


The obtained data allow to formulate the following important observations: 

• after the oxidation during 800 s at 1000 C, the El 10 cladding manufactured in accordance with the 
traditional method of Zr-l%Nb ingot preparation has typical indications of the breakaway oxidation; 

• the E110 standard cladding achieves the zero ductility threshold after the oxidation during somewhat 
longer than 800 s at this temperature; 

• the sponge types of the El 10 cladding oxidized at 1000 C during 2500 s have a significant margin of 
residual ductility, low hydrogen concentration in the prior P-phase, and the uniform corrosion type of 
oxidation; 

• the first demonstration of the breakaway oxidation of the sponge El 10 claddings appears after 5000 s of 
oxidation, hydrogen content in the cladding remains still low but the zero ductility threshold is achieved 
due to the decrease in the prior p-phase thickness and the increase of oxygen concentration in the clad¬ 
ding matrix. 


4.15 

























































Besides, the obtained data allow to formulate one new problem: for oxidation at 1000 C, the critical meas¬ 
ured ECRs corresponded with the zero ductility thresholds of standard and sponge ECR claddings are similar 
to both types of the cladding but the oxidation duration differs approximately six times. In the context of this 
problem, the following question could be formulated: is the measured (calculated) ECR a unequivocal crite¬ 
rion characterizing the zero ductility conditions? The additional consideration of this issue will be continued 
in one of next paragraphs of the report. To investigate the sponge El 10 cladding behavior at temperatures 
lower than 1000 C, one reference test was performed at 900 C (with modified steam generator). The results 
of this test are shown in Fig. 4.13. 


Cladding 

characterization 

Oxidation 

duration 

Cladding appearance 






E110 

(standard) 

ECR=6.7% 

t e f=4804 s 





#131 







El lOc^nj) 

ECR=7.5% 

t e t“ 14400 s 






#137 



i 



70 

60 

g 

.£» 

50 

o 

40 

3 

T3 

30 

03 

3 

rs 

20 

55 

<L> 

oc 

10 





o 









a E110, 900 C 

O El 10 G(3ru| , 900 C 

.E110, 1100C 



zr 

A 


I 

\ 








\ 

\ 

\o 








V 

A 











& 

» 

V 









r 

% 

i 









c 

\ 

> 

r* - «l 

b - 

- fr-ft-i 


e-900-en-v2 


6 8 10 12 
Measured ECR (%) 


14 


16 


18 


Fig. 4.13. The comparative test data characterizing the El 10 and E110 G (3ru) behavior 

after the oxidation at 900 C 


The present data allow to observe that general tendencies revealed for two types of the El 10 cladding at 

1000 C are retained at 900 C also: 

• the embrittlement of the E110 G (3ru) cladding is not the consequence of the breakaway oxidation in 
contrast to the standard El 10 cladding; 

• much more time is needed to achieve approximately the same ECR in the sponge El 10 cladding and the 
same margin of residual ductility as those in the standard El 10 cladding (14400 s and 4800 s 
respectively); 

• the critical measured ECR characterizing the zero ductility threshold of sponge type in the El 10 cladding 
(El 10 G (3ru)) at 900 C is approximately the same as that for the sponge El 10 at 1000 C and the standard 
El 10 at 1100 C (8.3% ECR) but the causes for embrittlement are different. The embrittlement of sponge 
El 10 is caused by the oxygen induced mechanism but the embrittlement behavior of the standard El 10 is 
determined by the combination of oxygen and hydrogen effects. 

The last position of temperature dependent investigations with the sponge El 10 cladding was devoted to the 

tests at 1200 C. The results of these tests are shown in Fig. 4.14. 


4.16 
















































ECR=7.8% 

(824 ppm of H content) 


ECR=22.3% 

(2200 ppm of H content) 



#120 



#112 


o 

~0 

"c3 

"v> 

o 

oc 


70 

60 

50 

40 

30 

20 

10 

0 


i 


8 


10 


12 


14 

ECR (%) 


16 


18 


20 










e 

1200-all-en-v2 


\ 




A 

E110 (sponge Zr), T=1200 C 

E110 (iodide/electrolytic Zr), 
T= 1000-1200 C 




\ 







\ 

\ 














\ 









_ 

\ 

1 









— 

- AA, 

Gk 

\ 

!—— 



r— — — 




■A- 


22 


24 


Fig. 4.14. The characterization of appearance and RT mechanical properties of the EllOcpru) cladding 

after the oxidation at 1200 C 


The major conclusion from these test is the following: for oxidation at 1200 C, and in spite of the presence of 
black bright oxide on the tested cladding, the zero ductility of sponge El 10 cladding of the El lOcoru) type 
was approximately the same as that for the standard El 10 cladding. The oxidation of the El lOooni) cladding 
corresponded to the significant hydrogen absorption. 

Thus, the results of these tests have demonstrated that the oxidation behavior of the El lOcom) cladding dete¬ 
riorated sharply at 1200 C. In this problem, it is interesting to note that ring compression tests performed at 
20 C in the ANL with Zry-4, Zirlo, M5 samples oxidized at 1200 C have shown that the residual ductility of 
these claddings “decreased rather abruptly from 5 to 10% ECR” [13]. These data confirm that the oxygen 
diffusion processes in zirconium claddings are considerably changed on proceeding from 1100C up to 
1200 C. However, the nature of these processes is not quite understood now. Besides, in spite of the similar¬ 
ity in the El 10 G (3ru) behavior and Zry-4, M5, Zirlo one at 1200 C, a significant difference has been revealed 
also. These differences are associated with the hydrogen content in the oxidized cladding. The embrittled 
Zry-4, Zirlo cladding have a low hydrogen content and (on our opinion) due to this reason these claddings 
have demonstrated the remarkable improvement in the cladding ductility at the temperature increase in me¬ 
chanical tests up to 135 C. The embrittled El 10 G (3ru) cladding has a high hydrogen content and the tempera¬ 
ture increase in mechanical tests up to 135 C has not led to the increase in residual ductility. In this connec¬ 
tion, it is reasonable to assume that the revealed difference in the behavior of these claddings is associated 
with the fact that El 10 o< 3 nj) is not 100% sponge material. 

4.3.4. The analysis of results obtained in the test with the E635 cladding fabri¬ 
cated using sponge Zr 

The E635 G( fr) cladding fabricated on the basis of 100% French sponge Zr was the last type of niobium¬ 
bearing claddings manufactured with the use of sponge Zr and tested in the frame of this work. These several 
tests were performed to extend the test data base developed to determine the sensitivity of the cladding oxi¬ 
dation behavior to the alloy chemical composition. The major results of tests with the E635 G(fr) cladding oxi¬ 
dized at 1100 C are presented in Fig. 4.15. 


4.17 







































ECR=12.5% 

(18 ppm of H content) 




70 


60 

o x 

50 



-i —' 

O 

3 

~o 

40 

3 

rs 

30 

C/5 

4) 

Cci 

20 


10 


0 


#100 


i 










T=1 100 C 
a standard 635 

□ E635 G(fr) 

-standard E110 


t 

i 

i 





— i- 

i 

t 

i 

* A 









» 

« 

1 

l 

A 




* A 





i 

» 

» 

A 1 , 

A 




A 

A 

i 

\ A A 

t 

□ 


e635-g635 


8 


12 


16 


20 


24 


ECR (%) 


Fig. 4.15. The appearance of the E635 C (f r ) cladding after the oxidation test at 1100 C and comparative 
E635 (standard), El 10 (standard), E635 C (f r ) results of ring compression tests 


The obtained data allow to note the following: 

• In spite of the fact that the E635 G (fr) oxidized cladding is characterized by a low hydrogen content in the 
prior [3-phase up to 12.5% ECR, the first visible indications of the breakaway oxidation were observed 
on the cladding surface at 11% ECR; 

• in contrast to sponge types of the El 10 cladding, the general difference in zero ductility thresholds of the 
standard E635 and sponge E635 was not revealed. 


4.3.5. The comparative consideration of a microchemical composition 
of different types of the El 10 alloy 

Results of tests performed with the sponge El 10 cladding have demonstrated a significant improvement in 
the El 10 corrosion resistance. To explain this phenomenon, the microchemical aspects of difference between 
the iodide/electrolytic El 10 and sponge El 10 alloys were considered. The first step in these investigations 
was connected with the analysis of results of previous investigations performed in this line. H.Chung 
developed a very interesting model for aliovalent elements specially for the interpretation of the revealed 
difference in the El 10 and M5 alloy behavior [12]. This model is based on Wagner electrochemical theory of 
oxygen ions passing through oxide to the oxide/metal interface with regard to the presence of impurity and 
secondary phase precipitates in the cladding material [15]. The employment of this theory performed by 
H. Chung using the aliovalent element model in our case allowed to develop the following conception of the 
oxide behavior at the high temperature oxidation (>800 C) of niobium-bearing alloys: 

• the impurity and alloying element should be subdivided into overvalent elements and undervalent 
elements in relation to the tetragonal zirconium; 

• the presence of overvalent elements leads to the decrease of O' vacancies and to the increase of the 
stoichiometric degree of oxide (a low fraction of the protective tetragonal oxide); 

• undervalent elements provide: 


4.18 






































- a high density of oxygen ions vacancies; 

- the tendency towards the retention of under-stoichiometric oxide with a high fraction of tetragonal 
protective oxide; 

• such binary alloys as the El 10 and M5 contain a definite quantity of undervalent beneficial impurities: 
Ca, Al, Mg, Fe, Ni and a definite quantity of one overvalent alloying element Nb. 

The comparative analysis of the M5 and standard El 10 fabrication processes including the surface finishing 
performed by H. Chung led him to the following conclusions: 

• both alloys have the same content of deleterious overvalent niobium; 

• the M5 alloy is enriched with such beneficial elements as Ca, Al, Mg, Fe, Y during the fabrication 
process; 

• the standard El 10 alloy is enriched with such specific deleterious impurity as F during the fabrication 
process (electrolytic Zr production); 

• the general difference in the oxidation behavior of the standard El 10 and M5 alloys is caused by the 
differences in the above listed beneficial and deleterious impurities. 

In the context of these conclusions, it should be specially noted that: 

• the use of fluoride compounds during the electrolytic El 10 fabrication was always the object of a special 
attention and control of F content during the El 10 manufacturing process; 

• the results obtained by V. Vrtilkova after the oxidation tests with the El 10 alloy were used by H. Chung 
on the validation of his position in this issue [16]. However, it should be pointed out that V. Vrtilkova 
used the El 10 tubes manufactured from iodide Zr only because only this method of the El 10 fabrication 
was employed in Russia before 1985. But it is known that the iodide Zr contains much less impurities 
than electrolytic and sponge Zr; 

• besides, Wagner theory could be used for the impurity elements having a significant solubility in ZrO: 
(Ca, Mg, Al, Y), but it is known that such elements as Fe, Nb, Sn are practically insoluble in Zr0 2 and 
these elements are present in the oxide on the crystalline grains as independent phases. 

Russian investigations performed on developing the El 10 alloy have shown that the most deleterious impuri¬ 
ties in the zirconium alloys are the following: C, N, F, Cl. Si. Besides, such elements as Ti. Al, Mo, Ta, V 
negatively influence the corrosion resistance, the beneficial influence was revealed for Fe and Cr. The influ¬ 
ence of major alloying elements such as Nb and Sn is not univocal. The corrosion resistance of zirconium 
alloys with the Nb and Sn alloying elements is the function of the content of these elements in alloys. The 
investigations in detail of the corrosion resistance sensitivity of Zr-Nb-based alloys and Zr-Sn-based alloys to 
the alloying element content performed recently confirmed this statement [9, 17]. The recent reassessment of 
results of autoclave corrosion tests (500 C was the maximum temperature) with the M5 cladding has shown 
that C (with the content >100 ppm), Al (with the content 20-150 ppm), N, Sn (with the content >100 ppm). 
Si (with the content >80 ppm) have a detrimental influence over the M5 corrosion behavior [7, 18]. In addi¬ 
tion to that: 

• the beneficial influence of Fe, Cr and Fe/Cr ratio over the corrosion resistance of Zr-Sn alloys was re¬ 
vealed [19, 20,21,22]; 

• the acceleration in the corrosion process was observed as a function of such impurities as Al, Ti, Mn. Pt. 
Ni, Cu in the Zr-2.5%Nb alloy [23]; 

• the optimal corrosion resistance was obtained for C and Fe impurities in the Zr-2.5%Nb alloy. It was 
revealed that the optimal content of these elements is about 30 ppm C and 1100 ppm Fe [23], 

In conclusion of this overview of previous investigations, the following information concerning such 
elements as Nb, O. Hf may be added: 


4.19 


• Russian investigations on the oxygen content in the Zr-Nb alloy of 400-1000 ppm and French investiga¬ 
tions on the oxygen content in the Zr-Nb alloy of 900-1800 ppm [18] have shown that the oxygen con¬ 
centration has the negligible influence over the corrosion resistance of niobium-bearing alloys; 

• the same investigation and results of investigations published in [9] allow to consider that the corrosion 
behavior of niobium-bearing alloys is not sensitive to the variations of Nb content in the alloy in the 
range 0.9-1.1 wt%; 

• the role of hafnium is not completely clarified now but this element is usually considered as neutral. 

Unfortunately, the comparative analysis in detail of this research results with results of previous 
investigations devoted to the relationship between the microchemical composition of niobium-bearing alloys 
and the corrosion behavior of these alloys allow to formulate the only following general conclusions: 

1. The high corrosion resistance of Zr-Nb alloy under operating conditions (a low temperature range) is the 
necessary requirement for this type of zirconium cladding but this requirement fulfillment does not guar¬ 
antee that the alloy will demonstrate a high corrosion resistance under LOCA relevant conditions (a high 
temperature range) also. 

2. In spite of the numerous investigations performed in this line, the nature of the relationship between the 
corrosion resistance and chemical composition of the alloy is not quite understood yet. 

Taking into account these conclusions, the plan for comparative studies of the chemical composition of the 
iodide/electrolytic El 10 and sponge El 10 was developed on the basis of the following major provisions: the 
measurement results of the standard El 10 alloy chemical composition (used for our oxidation tests) and 
measurement results of sponge types of the El 10 alloy chemical composition (tested during this program) 
must be compared. Reasonable differences in the content of chemical elements should be revealed and ana¬ 
lyzed. 

At the first stage of this plan, a typical content of the El 10 alloy presented in Table 4.4 was compared with a 
real content of chemical elements in each batch of tested El 10 alloy. This comparison showed that the typi¬ 
cal composition of the El 10 alloy was representative for the tested El 10 material except for Fe content. Real 
Fe content in the El 10 alloy used for this program was 86 ppm. 


Table 4.4. Chemical composition of the El 10 alloy (standard) [5] 


's 

Chemical element 

o. 

D. 

Al* 

B 

Be 

C 

Ca 

Cd 

Cl 

Cr 

Cu 

F 

Fe 

H 

Hf 

c 

a> 

<30 

<0.4 

<30 

<40-70 

<100 

<0.3 

<7 

<30 

<10 

<30 

140 

4-7 

300-400 

o 

o 


cd 

CJ 

’S* 

K 

Li 

Mn 

Mo 

N 

Ni 

O 

Pb 

Si 

Sn 

Ti 

Nb 

H 

<30 

<2 

<3 

<30 

<30-40 

<30-39 

500-700 

<50 

46-90 

<100 

<30 

o 

'sD 

Ln 

1 

1.1 wt% 


* Concentrations marked as <30,... reflect that fact that the real concentration of elements was not measured, because 
it was less than the low threshold of the detector (procedure) sensitivity 


At the second stage of the research, chemical compositions of the El l0 G(fr) , El 10 G(3fr) , El 10 G(3ru) alloys were 
compared with the standard El 10 composition. The comparative analysis of the appropriate data allowed to 
establish the reasonable difference in the impurity contents for Fe and Hf only. Nevertheless, other several 
beneficial and deleterious elements were added into the comparative results presented in Fig. 4.16. It should 
be noted that in spite of the fact that Hf influence over the corrosion resistance is not understood, we attrib¬ 
uted this element to deleterious impurities taking into account: 

- results of tests with the E110| OW H f cladding; 

- the fact that Hf stabilizes the monoclinic oxide at higher temperatures. 

By the way, the comparative analysis of the El 10| OwH f and standard El 10 cladding chemical compositions 
showed that the impurity contents in these materials were the same except for Hf content. 


4.20 




































Standard El 10 (average content in tested claddings) 

The real content is not known, but this content does not exceed appropriate limit 
Impurity range in tested El lOo,,,, El I0 rjl)lrl , El I0 C<3 „, 



c- •: 


r~ 

0 


100 200 300 400 

_Content of element (ppm) 


500 


Fig. 4.16. Comparison of some data on the impurity content in the standard El 10 and E110c<f rh 
E 110 Q 3 frh Ell 0 C ( 3 nj) at the beginning of the cladding fabrication 


Thus, the obtained data showed that one significant difference in the standard El 10 and sponge El 10 chemi¬ 
cal composition was observed: Fe content in the sponge El 10 is higher than that in the standard El 10. But 
the beneficial effect of iron on the corrosion behavior of zirconium alloys is known for a long time. There¬ 
fore. the majority of alloys including such Russian alloy as E635 have a high or quite high Fe content (see 
Table 4.5). In this case, such alloys as Zircalys. Zirlo and E635 employ iron as the alloying element. 


Table 4.5. Composition of zirconium alloys used in reactor fuel design [24] 


Element 

Zircalov-4 

ZIRLO 

E635 

M5 

E110 

Nb (wt%) 

— 

0.9-1.3 

0.95-1.05 

0.8-1.2 

0.95-1.05 

Sn (wt%) 

1.2-1.7 

0.9-1.2 

1.2-1.3 

— 

— 

Fe (wt%) 

0.18-0.24 

0.1 

0.34—0.40 

0.015-0.06 

0.006-0.012 

Cr (wt%) 

0.07-0.13 

— 

— 

— 

— 

Zr 

Balance 

Balance 

Balance 

Balance 

Balance 


The results obtained during this research confirm the important role of iron in the cladding oxidation behav¬ 
ior. This thesis is based on the fact that sponge types of the El 10 cladding with the higher content of iron 
demonstrate the better corrosion behavior than the standard El 10 alloy with low Fe content as well as on the 
fact that the results of oxidation tests of the E635 cladding (iodide electrolytic) with a very high content of 
iron have shown that the oxidation behavior of the standard E635 cladding is somewhat better than that of 
the standard El 10 cladding. Though, in this case the absence of a general difference in the oxidation behav¬ 
ior of the standard E635 and sponge E635 impels to involve additional test data to continue the analysis of 
this and other issues connected with the chemical composition of Zr-Nb alloys. The interesting scope of in¬ 
vestigations was recently performed in the VNIINM (Russia) [25]. 

Seven types of niobium-bearing claddings (iodide electrolytic) were oxidized at 1100C and mechanically 
tested. These seven types of the cladding are characterized by the following range of chemical composition 
variations: 

• Nb -> 0.9-11 wt%; 


4.21 



































• Fe -> 80-1400 ppm; 

• C —» 45-200 ppm; 

• Hf —> 100-430 ppm; 

• Cr —> 30-60 ppm; 

• O-» 350-1300 ppm. 

The analysis shows that these claddings have demonstrated quite different oxidation and mechanical behav¬ 
ior. These types of this behavior and appropriate contents of chemical elements in the cladding material are 
characterized in Table 4.6. 


Table 4.6. The organized results of oxidation and mechanical tests with seven types of the standard El 10 

and modified El 10 (the data of [25]) 


Test parameters 

Corrosion resistance at 10% ECR 

Best 

Intermediate 

Worst 

The cladding sample number 

##2, 3, 4 

##5,6 

##1, 7 

The corrosion type 

Uniform (lustrous 
black protective 
oxide) 

Beginning of the nodular 
oxidation (white spots on 
the cladding surface) 

Typical breakaway 
oxidation. The spallation 
of white oxide 

Hydrogen content (ppm) 

60-200 

200^100 

400-600 

Mechanical properties 

Maximum residual 
ductility and fracture 
energy 

Very low residual 
ductility and fracture 
energy 

Very low residual 
ductility and fracture 
energy 

The impurity content (ppm): 




Fe 

130-450 

80-1400 

100-140 

O 

350-600 

890-1300 

350-500 

C 

<45-65 

70-200 

60-80 

Hf 

100^30 

360-370 

100-360 

Cr 

<30-60 

<30 

<30 


The consideration of the data presented in Table 4.6 shows that: 

• the cladding with the iron content of about 130-450 ppm having relatively low C content and low Hf 
content (~100 ppm) have demonstrated the best corrosion resistance; 

• the intermediate corrosion resistance was observed at very low and very high Fe content (80 and 
1400 ppm) and high Hf content, in this case, the direct sensitivity to the C content in the range of 70- 
200 ppm was not revealed; 

• the relationship between the worst corrosion resistance and the concentration of chemical elements listed 
in Table 4.6 is not quite understood. 

The analysis of these results performed by the VNIINM researchers led them to the assumption that the oxi¬ 
dation behavior of the El 10 alloy is not so much the function of Fe, C, ... concentration as the function of 
the quantity of impurities in the alloy. Special measurements performed for this goal showed the following 
[25]: 

• the sum of Ni, Al, Si, Ca, K, F, Cl, Na, Mg impurities in the El 10 cladding, having the intermediate and 
worst corrosion behavior was about 110-135 ppm; 


4.22 












• the appropriate sum in the El 10 claddings that demonstrated the best corrosion behavior was about 25- 
45 ppm, in this case, the corrosion resistance decreased as impurity contents increased from 25 up to 
45 ppm. 

Unfortunately, it was impossible to perform the comparison of this type for impurity contents in the standard 
E110 and sponge El 10 on the basis of data presented in Table 4.4 and in the Fig. 4.16. The thing is that the 
methods used to measure the concentration of very many impurities allowed to guarantee that the concentra¬ 
tion of some impurity was not higher than the value presented in Table and Figure (<30, <40, ...). However, 
these methods did not allow to measure the real impurity concentration in the alloy. It was also impossible to 
estimate the impurity contents in the M5 alloy because these data were not published. But some additional 
data useful for the analysis of this issue could be taken into account from the consideration of impurity con¬ 
tents in Zirlo presented in Table 4.7. 


Table 4. 7. Chemical composition of the Zirlo cladding tube [26] 


Element 

content 

(ppm) 

Chemical element 

A1 

C 

Cr 

Cu 

Hf 

Fe 

Mg 

Mo 

Ni 

Nb 

N 

O 

Si 

Sn 

Ti 

w 

Zr 

120 

20 

10 

20 

<40 

1000 

<10 

<10 

<10 

1.23 

wt% 

50 

1450 

130 

1.08 

wt% 

19 

<40 

Ba¬ 

lan¬ 

ce 


A special comment must be done before the analysis of these data: 

• chemical compositions of the El 10 alloy presented in Table 4.4, Table 4.6 characterize the ingot compo¬ 
sitions; 

• the E110 chemical compositions presented in Fig. 4.16 characterize the beginning of the cladding fabri¬ 
cation processes; 

• special measurements of the El 10 overall impurity contents were performed in the unoxidized El 10 
tubes; 

• the Zirlo composition presented in Table 4.7 was measured using the Zirlo tubing. 

Thus, the cladding tubes were used in both cases for the measurement of the sum of the El 10 impurity con¬ 
tent and Zirlo impurity content. The comparison of appropriate data showed that the hypothesis proposed in 
[25] to explain the El 10 oxidation behavior on the basis of the sum of impurity contents was not confirmed 
as applied to the Zirlo claddings. So, the sum of two impurities only (A1 and Si) is 250 ppm in the Zirlo clad¬ 
ding. But the oxidation behavior of the Zirlo cladding is characterized by the uniform corrosion and low hy¬ 
drogen concentration in accordance with the data obtained in the ANL. Though, on the other hand, some data 
characterizing the content of impurities in the M5 alloy is in a good agreement with the assumption about the 
fact that the optimization of the content of beneficial impurities does not allow to achieve a high corrosion 
resistance at a high overall concentration of other impurities in the alloy. So, these some data characterizing 
the M5 alloy are the following [7, 18]: 

• the sum of Ca, Mg, Sn, S contents in the alloy is less than 1 ppm; 

• the sum of Si, Zn, A1 contents in the alloy is less than a few ppm; 

• the C content is 25-120 ppm. 

Besides, these data characterizing the chemical composition of the M5 alloy allow to return to the discussion 
of the fact that the electrochemical theory of the cladding oxidation and the model of aliovalent elements 
(developed on the basis of this theory) must be added with models taking into account other competing 
processes in the niobium-bearing claddings under high temperature oxidation conditions. 

If the experimental and analytical data presented in this paragraph are generalized then the following 
conclusions may be made: 


4.23 






















• the cladding high temperature oxidation behavior has a series of peculiar features associated with the 
phase compositions and temperature dependent characteristics of the solubility and diffusion of alloying 
and impurity elements in the oxide and metallic matrix. Therefore, the studies of these processes should 
be performed under high temperature conditions. A direct transfer of results for low temperature 
corrosion tests to the high temperature corrosion behavior may result in grand errors; 

• the impurity composition is one of the key factors determining the oxidation behavior of Zr-Nb alloys at 
high temperature conditions; 

• there is a serious reason to change the current approach from the classification of impurities by the 
beneficial and deleterious ones to the following approach: 

- minor alloying elements (this term was proposed in [23]), these are impurities allowing to provide the 
uniform mechanism of oxidation and to minimize the hydrogen content in the oxidized cladding; the 
requirements for the content of these elements must be developed; 

- deleterious impurities for which the requirements must be stipulated for their individual content in the 
alloy or the requirements for their total content in the alloy; 

- neutral impurities; the concentration of these impurities in the alloys in the current limits do not affect 
the oxidation mechanism and oxidation rate. 

As for the El 10 alloy, the list of impurities in accordance with this new classification cannot be determined 
basing on the results of this research. But it seems that iron is the first candidate for the incorporation into the 
set of the El 10 minor alloying elements. 

4.4. The comparative analysis of the El 10 material microstructure 

As it was reported at the beginning of this section, the bulk effects (in the context of the cladding oxidation 
behavior) are considered on the basis of two independent experimental data bases: 

1. The data base characterizing the dependence of the cladding oxidation behavior in a function of the clad¬ 
ding material chemical composition. 

2. The data base characterizing the dependence of the oxidation phenomena in a function of the cladding 
microstructure. 

In our case, the microstructure investigations with samples from different types of the El 10 cladding were 
especially urgent because of the fact that the analysis of results obtained in microchemical investigations 
presented in the previous paragraph did not allow to develop the unequivocal explanation of the El 10 
specific behavior. 

The program for special TEM examinations was worked out to determine the following comparative data 
characterizing the microstructure of tested claddings: 

• phase conditions and phase compositions; 

• the grain size in the cladding matrix; 

• the characterization of the secondary phase precipitates including: the chemical composition, size, 
density, the character of precipitates’ distribution. 

The following as-received tube samples were used for this research: 

• standard El 10; 

• E110 G( fr); 

• E110 low Hf- 

The basic examinations were performed in the RIAR using the transmission electron microscope JEM-2000 
FX II at the acceleration voltage 120 kV. 


4.24 


Taking into account that results of numerous investigations performed recently with different claddings 
showed a strong dependence ot the corrosion resistance on parameters of intcrmetallic precipitates in the 
alloy, the examinations in detail ot precipitates were carried out using thin foils and carbon replicas. 

The comparative data characterizing TEM images of different El 10 claddings are presented in Fig. 4.17 to¬ 
gether with the M5 TEM image reprinted from [27], The analysis of visual observations obtained after stud¬ 
ies of this information has shown that: 

• all presented cladding samples have the equiaxed a-Zr grains and globular secondary phase particles 
(SPP) uniformly distributed in the a-Zr matrix (see also Fig. 4.18); 


• the microstructure of all cladding samples is completely recrystallized. 



Fig. 4.17. Low magnification of TEM micrographs for El 10, E110 G(fr) , El 10| OVV n f claddings 

and the M5 cladding [reprinted from 27] 



Fig. 4.18. The characterization of the SPP distribution in the a-Zr grains of the El 10 

and El 10 G (fr)claddings 


4.25 










































The next stage of examinations was devoted to the determination of the a-Zr matrix chemical composition, 
the grain boundary and SPPs. The energy dispersion X-ray analysis (EDX) was employed for these 
quantitative measurements. Besides, the SPP sizes and SPP density were measured also. 

The measured chemical compositions of the a-Zr matrix and p-Nb precipitates in different El 10 claddings 
arc presented in Table 4.8. 


Table 4.8. Zr, Nb content in the matrix , grain boundary and J3-Nb precipitates 


Element 

Concentration (wt%) 

Matrix 

Grain boundary 

P-Nb 

E110 

E110[ OW Hf 

El 10 G (fr) 

El 10 

E 110| ow Hf 

E 1 10o(tr) 

E110 

E 110| OVV Hf 

El lOo(fr) 

Zr 

99.59 ±097 

99.34 ±103 

99.51*' 17 

99.54 ±lu 

99.62 ±092 

99 54± 114 

11.03 ±117 

9.66 ±048 

11,49 ±0 65 

Nb 

0.41 ±0 ' 26 

0.66 ±0 ' 28 

0.49 ±OJI 

0.46 ±0 ' 29 

0.38 ±0 ' 26 

0.46 ±0 ' 32 

88.97 ±2 ' 68 

90.34 1119 

88.51*' 56 


The analysis of obtained data showed that no significant differences regarding Zr and Nb content in the 
microstructure components of El 10, El 10 G (f r ), El 10| OwH f claddings were revealed. It is known that the best 
corrosion resistance is observed in the cladding material with fine a-Zr grains and fine P-Nb precipitates 
distributed uniformly. Therefore, the next task of TEM examinations was focused on measurements of SPP 
size distributions and on the determination of the average size for a-Zr grains and intermetallic precipitates. 

The results of appropriate measurements allowed to reveal the following: 

• the a-Zr average grain size in the tested cladding materials was very similar (2.8-3.2 pm); 


• the average size of secondary precipitates was similar also (43-48 nm), SPP size distributions in the 
El 10, El 1 0 G (fr), El 10| OW nr samples are presented in Fig. 4.19; 

• the SPP density was about 1.8x1O 20 m '. 

To improve the representativity of TEM data obtained in the RIAR with the use of a very limited number of 
samples, the independent TEM investigations were performed in another Russian scientific institute (Institute 
of Reactor Materials) also. The results of these examinations are in a good agreement with the RIAR data 
except for the SPP average size in the El 10 cladding which was determined as 60 nm. The SPP average size 
in El 10| OwH f and El 10 G ( fr) claddings were estimated as 55 and 41 nm, respectively. 



El 10 G( fr) sponge 
























)0 






















© x 

























- 




































2 

U- >2 


































6 - 
















= 

































10 2 

0 30 4 

0 30 6 

0 

] 

0 80 9 

Dian 

0 100 1 

ictcr 

0 1 . 

(n 

» 1 . 

Ill 

30 1* 

0 1 . 

50 160 1 

?0 If 

30 190 2 


Fig. 4.19. The SPP distribution in the El 10 and Ell 0 G (fr) claddings 


The analysis of TEM micrographs and TEM SPP distributions allowed to reveal the following general 
differences in the El 10 (El 10| OW Hf) and El 10 G( f r) cladding microstructures: 

• the E110 cladding contained only one type of secondary precipitates, this is a beta-phase particle 
enriched with niobium (86-91%); 


4.26 
































































































































• the EllOcxfr) cladding contained (in addition to (3-Nb precipitates) the intermetallic phase of the 
Zr(Nb.Fe) 2 type (see Fig. 4.20) with the average size 180 nm. 

To compare the obtained TEM results, the organized data base with parameters of iodide/electrolytic and 
sponge E110 was developed and presented in Table 4.9. Taking into account the limited number of TEM 
examinations pertormed in the frame of this work, the El 10 data were added with the results of VNIINM 
investigations [25]. Besides, the published data on the xM5 cladding [7] were incorporated in this Table also. 



- P- '? ^ _£_ D®® rus 

Fig. 4.20. High magnification of the SPP TEM micrograph in the El 10 G( fr) cladding 


Table 4.9. The comparative data characterizing the microstructure of El 10, El 10 G( f r) claddings 

and the \15 cladding [7'f 


List of parameters 

Cladding type 

E110 

(iodide/electrolytic) 

El IOgut) 

(sponge) 

M5 

(sponge) 

1. Intermediate and final 
annealing temperature (C) 

580 

580 

580 

2. The phase state due to 
the thermal treatment 

The fully recrystallized 
microstructure 
(aZr+pNb) 

The fully recrystallized 
microstructure 
(aZr+pNb) 

The fully recrystallized 
microstructure 
(aZr+pNb) 

3. The type of a-Zr grain 

Equiaxed 

Equiaxed 

Equiaxed 

4. The average size of a-Zr 
grain (pm) 

• 2.8 (this work) 

• 4-M.5 [25] 

3.2 

3-5 

5. The characteristics of 
P-Nb precipitates: 

• the geometrical form 

Globular 

Globular 

Globular 

• the average size (nm) 

45-60 (this work) 

50 [25] 

41—43 

45 

• the distribution density 
(cm' 3 ) 

1.8410' 4 

1.8-10 14 

1.510 14 

6. The intermetallic pre¬ 
cipitates: 





* The all data characterizing the M5 cladding was taken from [7] 

4.27 


























Cladding type 

List of parameters 

E110 

E 1 1 0o(fr) 

M5 


(iodide/electrolytic) 

(sponge) 

(sponge) 

• the type 

** 

Zr(Nb,Fe) 2 

Zr(Nb,Fe,Cr) 2 

• the size (nm) 

— 

180 

100-200 


The analysis of comparative data allows to conclude the following: 

• the most of the key parameters characterizing the cladding microstructure are practically the same in the 
El 10 (El lOiowHf), El 10 G <fr) and M5 claddings; 

• the major and the only revealed distinction between the El 10 and El 10| OW Hf claddings and the El 1 0 G( f r) 
and M5 claddings regards to the absence of iron-based precipitates in the cladding material. Some dis¬ 
tinction between the chemical composition of intermetallic SPPs in the Ell 0 G (fr) and the M5 alloy 
(Zr(Nb,Fe) 2 and Zr(Nb,Fe,Cr) 2 , respectively) may be associated with an insufficient scope of appropriate 
measurements performed in the El 10 G (fr) samples. 

The revealed differences in the presence and absence of iron-based precipitates are in the complete agree¬ 
ment with the results of microchemical investigations presented in the previous paragraph and with the data 
characterizing the iron solubility limit in zirconium claddings. This limit is about 100 ppm. Taking into ac¬ 
count that iron content in the El 10 alloy is about 90 ppm, it is fully dissolved in the zirconium matrix. Iron 
concentration in the sponge types of the El 10 cladding is in range 130-430 ppm. Therefore, one part of iron 
is dissolved in the matrix and the second part is presented as the intermetallic precipitates. As it was men¬ 
tioned in the previous sections of the report, the most of the appropriate investigations have demonstrated a 
high importance of intermetallic precipitates in the oxidation behavior of zirconium cladding. Moreover, it 
was revealed that the iron-based precipitates improve significantly the corrosion resistance and reduce H 
uptake. 

But the analysis of this research data and data obtained in the VNIINM [25] allows to assume that there is an 
optimal content of iron in the alloy with which the best corrosion behavior is observed. Lower and higher Fe 
concentrations deteriorate the corrosion resistance. 

The following general conclusions may be made on the basis of the whole scope of obtained results: 

• the differences in the oxidation behavior of the standard El 10 and sponge-based El 10 claddings are not 
a function of the cladding fabrication; 

• the differences in microchemical compositions of impurities are probably the major factor for the 
different oxidation behavior of these types of the El 10 claddings; 

• more careful measurements of the chemical composition of impurities especially such as C, N, F and 
other nonmetallic deleterious impurities must be performed in the future to adjust these phenomena; 

• additional investigations must be performed also to adjust Hf and Cr influence; 

• special experimental investigations with the El 10 cladding containing different Fe contents may also be 
useful for the determination of the optimal Fe concentration. 

4.5. The oxidation kinetics of the sponge type El 10 cladding 

The oxidation tests with the El 10 claddings manufactured on the basis of sponge Zr showed that the oxida¬ 
tion kinetics for the standard iodide/electrolytic and sponge El 10 cladding was similar, though some ten¬ 
dency towards the increase of oxidation rate was observed for the sponge type El 10 cladding in the tempera¬ 
ture range 1100-1200 C (see Fig. 4.21). These results are in the good agreement with the published data 
characterizing the oxidation kinetics of the M5 cladding [18]. 


* The iron-based intermetallic precipitates was not observed in the El 10 cladding 


4.28 














Unexpected results on the oxidation kinetics of the sponge El 10 were obtained in the temperature range 
900-1000 C. In accordance with experimental results presented in Fig. 4.22, the oxidation rate of the sponge 
E110 was much less than that of the standard El 10. The comparison of the sponge El 10 behavior with the 
M5 behavior demonstrated the same effect for this alloy also. 

The processing of data base characterizing the oxidation kinetics of the El 10c< fr ,, El 10 G( 3 fri, El IOgo^ clad¬ 
dings allowed to develop the approximation for the oxidation rate of the sponge type El 10 claddings as a 
function of the reciprocal oxidation temperature. The comparative data on the oxidation rate of two types of 
the E110 cladding are presented in Fig. 4.23. 




Fig. 4.21. The oxidation kinetics of different types of Zr-l%Nb claddings at 1100 and 1200 C 


4.29 
























20 


E 

o 

'bO 

c- 

c 


c3 

bO 

.£= 

bp 

’5 

£ 


16- 


12 


O' 








- 

900 C 

A Ell 0 G(3ru) 

E110 (iodide/electrolytic), 
this work 











. - * 

«« — 

— — 

*** 

***** * 

— 


A 

. -- 

/ 


— 

g-900 


0 


4000 


8000 
Time (s) 


12000 


16000 



Fig. 4.22. The oxidation kinetics of different types of Zr-l%Nb claddings at 900 and 1000 C 


4.30 













































10 


CJ 

rj 

SO 


"cS 

s— 

c 

c 

ca 

72 

>< 

C 


0.1 7 


0.01 


0.001 


0.0001 


standard El 10, this work 
E110, 


, G(3ru,3 fr.fr) 


E110 


G(3ru.3 fr.fr) 


K: 





x 

\ -V 

\ X 
\ 

\ 





\ \ 

\ X 

\ \ 

\ 

\ 

\+ 





k-l-g 


0.0006 


0.001 


0.0007 0.0008 0.0009 

1/T (K') 

Fig. 4.23. The characterization of the oxidation rate for the standard El 10 (iodide/electrolytic) 
and E110 claddings manufactured with the use of sponge Zr in the temperature range 900-1200 C 


This effect was adjusted basing on the analysis of the following microstructural characteristics of the 
oxidized cladding: 

• the comparison of ZrCf and a-Zr(O) thicknesses formed in the iodide/electrolytic and sponge El 10 dur¬ 
ing the oxidation at 1000 C (Fig. 4.24); 

• the comparison of ZrCf and oi-Zr(O) thicknesses formed in the sponge type El 10 cladding during the 
oxidation at 1000 and 1100 C (Fig. 4.25). 

The procedure of these studies was complicated by the fact that metallographic samples prepared from the 
standard El 10 claddings oxidized at 1000 C are characterized by a partial or complete loss of ZrCF layer. 
The oxide flake off occurred during the oxidation post-oxidation manipulations (including cutting of the 
oxidized cladding on preparing the metallographic sample). Nevertheless, the consideration of appropriate 
metallographic samples allowed to select the fragment of the polished sample with the representative thick¬ 
ness of ZrO: and the fragment of the etched sample for the estimation of a-Zr(O) thickness (see Fig. 4.24). 

The analysis of this data allowed to reveal a general difference in two main competing processes, which de¬ 
fine the oxygen uptake by the zirconium cladding: oxygen uptake during the oxidation of a relatively narrow- 
surface layer still to ZrCF and the oxygen uptake due to the oxygen transport into the cladding depth with the 
formation of a-Zr(O) layer. The comparative data presented in Fig. 4.24 show that ZrO: is thicker and 
a-Zr(O) is thinner in the iodide/electrolytic El 10 than those in the sponge El 10 at approximately the same 
weight gain. The similar case is observed at the comparison of the sponge El 10 oxidized at 1100 C with the 
sponge E110 oxidized at 1000 C (see Fig. 4.25). 

It is obvious that this specific behavior of the sponge type zirconium-niobium binary alloys in the tempera¬ 
ture range 900-1000 C is a function of the behavior of minor alloying elements (impurity elements) in this 
temperature range. The clarification of the physical nature in these processes is the task for future investiga¬ 
tions. 

Moreover, the extension of experimental data base, which will be obtained due to future investigations will 
allow to develop the practical approach to the assessment of the safety criteria for this temperature range 
because two opposite effects must be taken into account: 

1. Decrease of the oxidation rate. 


4.31 















2. Increase of the brittle a-Zr(O) layer thickness and, consequently, the reduction of the ductile prior 
P-layer after comparing as a function of time (taking into account the phenomena associated with p. 1). 


Sponge type of £110 


lodidc/cicctrohtic E110 


#98-4 


#44-5 


ECR-6.9% 


ECR=7.7% 

Etched 


Polished 

Etched 



Fig. 4.24. The difference in the thicknesses of Zr0 2 and a-Zr(O) layers in the El 10 cladding 
of sponge and iodide/electrolytic types at the oxidation at 1000 C 


#101-4 


#89-4 




1100 c 


ECR=10.5% 



1000 C 
ECR=8.9% 


<-ZrO,-» 


<-a-ZrO 


Fig. 4.25. The difference in the formation of oxide and a-Zr(O) layers in the El 10 cladding 

of sponge type at 1000 and 1100 C 


The preliminary recommendations concerning this issue will be presented in the Summary of the report. 


4.32 

































































































References for Section 4 


[1] Brachet J., Pelchat J., Hamon D., Maury R., Jaques P., Mardon J. "Mechanical Behavior at Room Tem¬ 

perature and Metallurgical Study of Low-Tin Zry-4 and M5™ (Zr-NbO) Alloys after Oxidation at 
1100 C C and Quenching", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Tran¬ 
sient andLOCA Conditions". Halden, Norway, September 10-14, 2001. 

[2] Mardon J., Frichet A., LeBourhis A. "Behavior of M5™ Alloy under Normal and Accident Conditions", 

Proc. of Top Fuel-2001 Meet.. Stockholm, May 27-30. 2001. 

[3] Ageenkova L., Braslavsky G., Kishenevsky V., Nikulina A., Tararaeva E., Shebalov P., Zaimovsky A 
“Structural and Phase Transformation in Zr-Nb Binary Alloys and their Relation to Properties”, Proc. of 
Conference on Reactor Physical Metallurgy, v.6, 29.5-1.6 1978, Alushta. Russia (rus). 

[4] Zaimovsky A., Nikulina A., Reshetnikov N. “Zirconium alloys in the Nuclear Industry”, Moscow, Ener- 

goizdat, 1981 (rus). 

[5] Shebaldov P., Peregud M., Nikulina A., Bibilashvili Yu., Lositski A., Kuz'menko N., Belov V., No¬ 
voselov A. “E110 Alloy Cladding Tube Properties and their Interrection with Alloy Structure Phase 
Conditions and Impurity Content”, Proc. of Zirconium in the Nuclear Industry': Twelfth International 
Symposium, ASTM STP 1354 (p. 545), 2000. 

[6] Reshetnikov F., Bibilashvili Yu., Golovnin I. “The Development. Manufacture and Operation of Fuel 

Rods of Nuclear Power Reactors”, Moscow, Energoatomizdat (rus), 1995. 

[7] Mardon J., Charquet D., Senevat J. “Influence of Composition and Fabrication Process on Out-of-Pile 

and In-Pile Properties of M5 alloy”, Proc. of Zirconium in the Nuclear Industry': Twelfth International 
Symposium, ASTM STP 1354 (p. 505), 2000. 

[8] Jeong Y.. Kim H., Kim T., Jung Y. “Effect of Nb-concentration, Precipitate, and Beta Phase on Corrosion 

and Oxide Characteristics of Zr-xNb Binary Alloys”, Zirconium in the Nuclear Industry: Thirteenth In¬ 
ternational Symposium (poster paper). 

[9] Jeong Y.. Kim H., Kim T. “Effect of (3-phase, Precipitate, and Nb-concentration in Matrix on Corrosion 

and Oxide Characteristics of Zr-xNb Alloys”, Journal of Nuclear Materials , v.317, 2003. 

[10] Kalin B. “The Investigation of Structure of Zirconium Alloys at the Alloy Modifications with the Use of 
Power and Radiation Attacks”. Proc. of Nuclear Power on the Threshold of the XXI Century' Confer¬ 
ence, 8-10 June 2000, Electrostal, Russia (rus). 

[11] Kalin B., Osipov V., Volkov N, Gurovich B.. Atalikova I., Menaykin S. “The oxide morphology in the 
zirconium alloys after its modification by the ion alloying technic”, Proc. of the 5 th Conference on the 
Reactor Physical Metallurgy', Dimitrovgrad. Russia. 8-12 September 1997. 

[12] Chung H.M. “The Effects of Aliovalent Elements on Nodular Oxidation of Zr-based Alloys”, Proc. of 
the 2003 Nuclear Safetx Research Conference , Washington DC, October 20-22, 2003, NUREG CP- 
0185,2004. 

[13] M. Billone, Y. Yan and T. Burtseva. “Post-Quench Ductility of Advance Alloy Cladding”, NSRC Con¬ 
ference, Washington DC, USA. October 20-22, 2004. 

[14] Lunde L., Videm K. “Effect of Material and Environmental Variables on Localized Corrosion of zirco¬ 
nium Alloys”, Proc. of Zirconium in the Nuclear Industry' Conference, ASTM STP 681, 1979. 

[15] Wagner C., Naturwissenschaften, 31(1943), 265. 

[16] Vrtilkova V., Valach M„ Molin L. "Oxiding and Flydrating Properties of Zr-l%Nb Cladding Material in 
Comparison with Zircaloys", Proc. of IAEA Technical Committee Meeting on "Influence of Water 
Chemistry on Fuel Cladding Behavior ", Rez (Czech Republic), October 4-8, 1993. 

[17] Takeda K.. Anada FI. “Mechanism of Corrosion Rate Degradation Due to Tin'’, Proc. of Twelfth Inter¬ 
national Svmposium “ Zirconium in the Nuclear Industry ", ASTM STM 1354 (p. 592), 2000. 

[18] Mardon J.P., Waeckel N. “Behavior of M5™ Alloy under LOCA Conditions”, Proc. of Top Fuel-2003 
meetings “ Nuclear Fuel for Today and Tomorrow Experience and Outlook", March 16-19, 2003, 
Wurzburg, Germany. 


4.33 


[19] Charquet D. “Microstructure and Properties of Zirconium Alloys in the Absence of Irradiation”, Proc. of 
Twelfth International Symposium: Zirconium in the Nuclear Industry, ASTM STP 1354. 

[20] Kakiuch K., Itagaki, Furuya T., Miyazaki A., Ishii Y., Suzuki S., Yamawaki M. “Effect of Iron for Hy¬ 
drogen Absorption Properties of Zirconium Alloys”, Proc. of Top Fuel-2003 Meeting “Nuclear Fuel for 
Today and Tomorrow Experience and Outlook”, March 16-19, 2003, Wurzburg, Germany. 

[21] Murai T., Isobe T., Takizawa Y., Mae Y. “Fundamental Study on the Corrosion Mechanism of Zr-0.2Fe, 
Zr-0.2Cr, and Zr-0.1Fe-0.1Cr Alloys”, Proc. of Twelfth International Symposium: Zirconium in the Nu¬ 
clear Industry, ASTM STP 1354. 

[22] Warr B.D., Perovic V., Lin Y.P., Wallace A.C. “Role of Microchemistry and Microstructure on Vari¬ 
ability in Corrosion and Deuterium Uptake of Zr-2.5Nb Pressure Tube Material”, Proc. of Zirconium in 
the Nuclear Industry: Thirteenth International Symposium, ASTM STP 1423. 

[23] Ploc R.A. “The Effect of Minor Alloying Elements on Oxidation and Hydrogen Pickup in Zr-2.5Nb”, 
Proc. of Zirconium in the Nuclear Industry: Thirteenth International Symposium, ASTM STP 1423. 

[24] Meyer R. “NRC Activities Related to High Bumup New Cladding Types, and Mixed-Oxide Fuel”, 
Proc. of an International Topical Meeting on Light-Water-Reactor-Fuel-Performance, April 10-13, 
2000, Park City, Utah, USA, (v.l). 

[25] Nikulina A.V., Andreeva-Andrievskaya L.N., Shishov V.N., Pimenov Yu.V. “Influence of Chemical 
Composition of Nb Containing Zr Alloy Cladding Tubes on Embrittlement under Conditions Simulating 
Design Basis LOCA”, Fourteenth International Symposium on Zirconium in the Nuclear Industry’, 
June 13-17, 2004. 

[26] Motta A., Ervin K., Delaire O., Birtcher R. Chu Y., Maser J., Mancini D., Lai B. “Synchrotron Radia¬ 
tion Study of Second-Phase Particles and Alloying Elements in Zirconium Alloys”, Proc. of Zirconium 
in the Nuclear Industry: Thirteenth International Symposium, ASTM STP 1423. 

[27] D.Gilbon et.al. "Irradiation Creep and Growth Behavior, and Microstructural Evolution of Advanced Zr- 
Base Alloys", Proc. of Zirconium in the Nuclear Industry: Twelfth International Symposium, ASTM- 
STP 1354. 


4.34 


5. Summary 


During 2001-2004, research was performed to develop test data on the embrittlement of niobium-bearing 
cladding of the WER type under LOCA relevant conditions. The program variability is shown in Fig. 5.1. 


Type of cladding material 


Un irradiated Irradiated 

~ I V ^ '-v 1 T : 

E110 E110A E110K EllOp EllOrn E110 c E635 E635 0 i Zr>-4 El 10 


Test conditions 


Oxidation t> pe 


Sinizlc-sidcd 


Double-sided 


Heating and 
cooling rates 
combination 


Oxidation 

temperature 






F/F 


F/Q 


F/S S/S 


S/F 




8(H) C 

9(H) C 


950 C 


1000 C 


1100 c 


1200 C 


Type of 
mechanical test 




Ring 

compression 


Ring Three-point 

tensile bending 


Temperature 
of mechanical 
test 


r -- 

20C 135 C 


200 C 


100 c 


Fig. 5.1. Outline of the research program 


The results of previous investigations suggested the following list of tasks for the program first part: 

• procedure development and validation to determine the zero ductility threshold; 

• determination of the zero ductility threshold sensitivity to transient conditions of the LOCA scenario 
(variations of heating and cooling rates during the oxidation); 

• development of the experimental data base characterizing the oxidation kinetics in the temperature range 
800-1200 C and zero ductility threshold of Zr-l%Nb cladding of the WER type (the El 10 alloy); 

• comparative analysis of the oxidation and mechanical behavior of Zircaloy-4, El 10 (Zr-l%Nb) and other 
niobium-bearing claddings. 


5. 1. Major findings of the program first part 


5.1.1. Methodological aspects of mechanical tests 

Consideration of the safety problem associated with the prevention of a fuel rod fragmentation caused by the 
embrittlement of a fuel rod cladding during the high temperature oxidation under LOCA conditions has 
shown that two general approaches could be used to estimate the appropriate phenomena: 

1. An approach based on the determination of the oxidized cladding fragmentation threshold. 

2. An approach based on the determination of the cladding material embrittlement threshold. 

5.1 
























































































From the formal point of view, the approach based on the determination of the cladding fragmentation 
threshold is preferable because the loss of resistance to the combination of accident loads is directly 
estimated in this case. Impact tests and thermal-shock tests are typical examples of this approach practical 
implementation. However, to use results of these tests for the safety analysis, the representativity of test 
conditions should be demonstrated in comparison with the combination of loads under real accident 
conditions. But it is known that real loads on the embrittled cladding during the accident cannot be estimated 
at present with sufficient accuracy. Besides, the whole previous experience shows that the cladding 
fragmentation threshold cannot be less than the cladding embrittlement threshold. Moreover, taking into 
account that embrittlement threshold is the basic material property this threshold depends only weakly on the 
type of mechanical tests. Therefore, the conservative approach was used to estimate the Zr-l%Nb (El 10) 
embrittlement condition. 

Special scoping tests were performed to develop a comparative data base characterizing the El 10 zero 
ductility threshold using the following types of mechanical tests: 

• ring tensile; 

• ring compression; 

• three-point bending. 

In accordance with the obtained data, the ring tensile and ring compression tests led to the same zero ductility 
threshold of the oxidized cladding. Three-point bending tests overestimated this threshold in comparison 
with ring compression and ring tensile tests. 

Taking into account these results and the fact that ring compression tests were used for the validation of the 
current safety criteria for the zircaloy cladding, this type of mechanical tests was chosen for the research. 
Some characterization of the oxidized cladding behavior under these test conditions are presented in Fig. 5.2. 


Load 



\ \ \ \Lipta-jour\\E \ Wegorova \ \ WRSM-2002\ \ Figures'^ \I J -scheme-ring- tests-v2. tif 



Displacement 


Fig. 5.2. The schematics of ring compression tests 


The additional analysis has shown that in spite of the fact that ring compression tests have the thirty year 
history many issues concerning the procedure of these tests still remain to be solved. Thus, the previous 
popular approach to the processing of ring compression test results based on the determination of relative 
displacement at failure (the sum of “elastic” and “plastic” displacement divided by the cladding outer diame¬ 
ter) did not allow to estimate the zero ductility threshold in the explicit form (see Fig. 5.3). Besides, the 
length of ring samples varied in the range 6-30 mm and the procedure for the fracture determination on the 
basis of load-displacement diagrams was not validated. 


5.2 















































Ring compression tests results 


E 

\ / 


CJ 

rs 

i / 

\f 
» / 

f 

t 'l iterion of ductile-to-hrittle 

r' 

transition: 

-T. 


must he determined on 

O 

> 


the basis of other types 



of investigations 


\ 

Zero ductility threshold 



Equivalent Cladding Reacted 


Fig. 5.3. The schematic of previous approach to the determination of zero ductility threshold 

Using results of a special experimental subprogram the following outcomes were obtained: 

• an approach to the determination of the zero ductility threshold on the basis of processing of load- 
displacement diagrams has been selected; 

• major provisions of this approach are based on the direct connection between the ductility margin and 
plastic strain value. To estimate the plastic strain, the parameter named the residual ductility at failure 
has been defined (see Fig. 5.4). When the residual ductility at failure tends to zero the ductile-to-brittle 
transition (zero ductility threshold) occurs in the oxidized cladding; 

• it is obvious that the selected method allows to obtain the macroscopic estimation of the zero ductility 
threshold only. But the fractography examinations performed with several samples confirm that the 
macroscopic and microscopic experimental data are in a good agreement; 

• to determine the fracture condition (the first through-wall crack) at the processing of load-displacement 
diagrams, a special reference data base has been developed; 

• the ring compression tests performed with rings of different length have shown that the residual ductility 
at failure does not depend on ring length in the range 8-25 mm for specially prepared ring samples. This 
special preparation consists in the elimination of the oxidized sample end parts (5 mm approximately) 
before the ring compression test. In other case the zero ductility threshold may be underestimated 
significantly. 



Fig. 5.4. The interpretation of ring compression test results on the basis 

of the load-displacement diagram 


5.3 













5.1.2. Methodological aspects of oxidation tests 


A special scope of analytical and experimental work was performed to validate the following parameters ot 
the oxidation facility: 

• the uniformity of the temperature distribution along the long cladding sample (100 mm); 


• the representativity of the cladding temperature measurements; 


• the absence of steam starvation conditions around the cladding sample. 

Taking into account the likely spallation and loss of zirconium dioxide on the oxidation of the El 10 cladding 
a special method for the weight gain measurement was developed, verified and implemented. 

To determine the zero ductility threshold sensitivity to such parameters of the oxidation scenario as heating 
and cooling rates, special scoping tests were performed. The major parameters of these tests are shown in 
Fig. 5.5. 


<D 


C3 

S_ 

<D 

Oh 

£ 

s 





Time 


Fig. 5.5. The variability of oxidation tests with different heating and cooling rates 


Numerous oxidation tests were performed with five combinations of heating and cooling rates: 
S(0.5 C/s)/F(25 C/s), S(0.5 C/s)/S(0.5 C/s), F(25 C/s)/S(0.5 C/s), F(25 C/s)/F(25 C/s), F(25 C/s)/ Q(200 C/s) 
where S means slow, F means fast, and Q means quench. Test results have shown that: 

• the specific features of the El 10 oxidation behavior are present at any combination of heating and 
cooling rates; 

• the most pronounced negative phenomena accompanying the El 10 high temperature oxidation are 
observed at slow heating rate (0.5 C/s). 

Nevertheless, the comparative analysis of zero ductility thresholds obtained for these five combinations of 
heating and cooling rates allowed to make the following important conclusions: 

• the zero ductility threshold has a low sensitivity to the combination of heating and cooling rates; 

• the zero ductility threshold of the oxidized cladding at cooling with 25 C/s (F) and 200 C/s (Q) is 
practically the same. 

These test results allowed to perform all subsequent investigations using the F/F (25 C/s/25 C/s) combination 
of heating and cooling rates. 

In the frame of this research program, almost all oxidation tests were performed under double-sided 
oxidation conditions. But to widen the data base for the comparative analysis, a special subprogram was 
carried out under single-sided oxidation conditions. The obtained results have shown that the zero ductility 
threshold of the El 10 is increased from about 8% ECR for double-sided oxidation to about 11% for single¬ 
sided oxidation. 

The general verification of experimental procedures developed for this program was the final stage of 
methodological work. The verification was performed in reference tests with Zry-4 cladding. After that, the 
obtained results were compared with the numerous published data on this alloy. Results of this comparison 


5.4 










have shown that the set of experimental procedures developed to determine the zero ductility threshold of Zr- 
l%Nb (El 10) cladding do not introduce large systematic errors. 


5.1.3. The embrittlement behavior of Zr-l%Nb (El 10) cladding 


This stage of the research program was performed with the use of commercial as-received El 10 cladding 
tubes. The double-sided oxidation tests w r ere carried out with these tubes in the temperature range 800- 
1200 C at F/F (25 C/s/25 C/s) and F/Q (25 C/s/200 C/s) combinations of heating and cooling rates. The 
embrittlement characteristics of oxidized claddings were determined using the ring compression tests at 20 C. 
For reference, mechanical tests were also performed at 135. 200. 300 C. This data base was supplemented 
with results of reference tests w ith the Zry-4 cladding, which was oxidized at 1100 C and tested at 20 and 
135 C. 

The visual observations of oxidized El 10 claddings have shown that this material is characterized by the 
initiation of the breakaway oxidation and earlier embrittlement in comparison with the Zry-4 cladding. So, 
some experimental data obtained at 1100 C and presented in Fig. 5.6 allow to make the following general 
conclusions (all ECR values are as measured): 

• the uniform oxidation mode characterizes the El 10 oxidation at low ECR (0-6.5% in this case). This 
oxidation mode leads to the formation of a black protective oxide on the cladding surface; 

• the breakaway oxidation mode occurs with the increase of the ECR up to 10.5% (in presented case). This 
oxidation mode is accompanied by oxide spallation; 

• the reference Zry-4 cladding surface oxidized at 11.3% ECR is covered with the black lustrous oxide. 
This fact confirms that the breakaway oxidation condition w as not achieved in the Zry-4 cladding mate¬ 
rial. 


Cladding 

material 


As-measured 
ECR (%) 


E110 


Zry-4 


6.5 


10.5 


11.3 


Appearance of 100 mm oxidized samples 




Fig. 5.6. The appearance of the El 10 and Zry-4 claddings (1100 C) as a function of the ECR 

The reassessment of previous investigations allow to establish that the cladding embrittlement after the 

breakaway oxidation is a sum of two processes: 

1. The oxygen induced embrittlement caused by the formation of ZrO : and a-Zr(O) brittle layers, the de¬ 
crease of the prior P-phase thickness and the increase of oxygen concentration in the prior P-phase. 

2. The hydrogen induced embrittlement caused by the hydrogen absorption in the prior P-phase and 
hydriding of the cladding material. 

Taking into account these considerations, the data base characterizing the residual ductility of the oxidized 

E110 cladding as a function of the ECR was supplemented with the numerous measurements of hydrogen 

concentration in the oxidized claddings (see Fig. 5.7). 

The major outcomes of results obtained at 1100 C are the following: 

• the El 10 oxidized cladding has a very high ductility margin and low' hydrogen content in the cladding 
material at the ECR up to 7.0%; 
















• a sharp decrease of the El 10 ductility occurs in the range of 7-8% ECR. This process corresponds to a 
sharp increase of hydrogen content in the prior [3-phase of the El 10 oxidized cladding; 

• the E110 zero ductility threshold is achieved at 8.3% ECR (as-measured) with 300-400 ppm of hydrogen 
content; 

• the Zry-4 reference cladding has demonstrated sufficient margin of residual ductility and very low hy¬ 
drogen content at 11.3% ECR. 



Fig. 5.7. The E110 residual ductility and hydrogen concentration as a function of the ECR after the 
double-sided oxidation at 1100 C and F/F, F/Q combinations of heating and cooling rates 


To understand and to interpret the obtained results, special post-test examinations of metallographic samples 
were performed on the basis of the optical microscopy, SEM investigations. Auger spectroscopy, 
fractography, and microhardness measurements. The analysis of these results with the support of data 
obtained during previous investigations allowed to make the following important observations and 
comments: 

• the cracking and spallation of the Zr0 2 layer observed on the El 10 cladding at the ECR higher than 7% 
provide hydrogen penetration into the oxide metal interface; 

• the cracking and spallation of the Zr0 2 layer is the result of transition from the understoichiometric 
protective tetragonal oxide to the stoichiometric porous monoclinic oxide; 

• this tetragonal-monoclinic phase transition is a function of temperature and volume stresses; 

• the stabilization of the protective tetragonal oxide prevents the hydrogen penetration into the cladding 
material; 

• most impurities in Zr0 2 that oxidize slower than Zr will stabilize the tetragonal oxide. It appears that a 
minor concentration of such elements as Sn, Fe, Cr in the zirconium dioxide allows to achieve this effect. 
This observation is in a good agreement with the Zry-4 alloying composition and oxidation behavior; 

• but the presence of some other precipitates leads to the formation of heterogeneous Zr0 2 characterized 
by high volume stresses and the tendency towards early tetragonal-monoclinic transformation. It is 
obvious that Zr0 2 on the El 10 cladding surface contains this type of precipitates. 


5.6 



































The analysis of the wavelength dispersive x-ray (WDX) dot maps and electron probe microanalyzer (EPMA) 
results obtained in the frame of this work has shown that the redistribution of niobium in the a-Zr(O) layer is 
observed in the Zr-l%Nb oxidized cladding. This process is characterized by the formation of radially ori¬ 
ented Nb-enriched and Nb-depleted areas. This phenomenon may facilitate the development of the heteroge¬ 
neous oxide (Zr 02 (Nb 205 )) with the tendency towards spallation. Besides, taking into account that niobium 
stabilizes the (3-Zr phase, the transformation of niobium enriched areas into the a-Zr(O) phase occurs at a 
higher oxygen concentration than in the areas with the lower niobium concentration. This effect is responsi¬ 
ble for the irregular boundary front between a-Zr(O) and prior [3-phase in these alloys. 

In spite of the fact that the El 10 oxidation rate is somewhat less than that in the Zry-4, the El 10 a-Zr(O) 
thickness is larger due to the difference in the allotropic transition temperature of the a-Zr(O) phase in the 
E110 alloy produced at the higher oxygen concentration in the (3-phase. This fact and the irregular 
a-Zr(0)/p-phase boundary lead to the reduction in the effective prior (3-phase thickness and to the increase of 
oxygen concentration in the prior [3-phase. The tendency towards the uniform oxygen distribution across the 
prior P-phase does not improve the ductility margin also. But the analysis of the microhardness measure¬ 
ments has shown that the revealed differences in the embrittlement behavior of the El 10 and Zry-4 claddings 
cannot be explained by the differences in the oxygen concentration in the P-phase because the El 10 cladding 
with the high ductility margin and El 10 cladding on the zero ductility threshold are characterized by similar 
microhardness values. This result confirms that the hydrogen absorption by the El 10 cladding is the key 
factor determining the El 10 embrittlement behavior. The hydrogen diffuses in the prior P-phase along ra¬ 
dially oriented a-Zr(O) grains, the Nb-enriched P-phase lines in the a-Zr(O) are the channels for the hydro¬ 
gen penetration also. Taking into account that the hydrogen solubility in the zirconium matrix is very sensi¬ 
tive to the temperature and, besides, the ductility of solid hydrides is strongly dependent on temperature, the 
mechanical behavior of the El 10 oxidized cladding was compared at 20 C and 135 C (the coolant saturation 
temperature at the end of the LOCA reflood mode). The obtained results have shown that: 

• in the case when the whole absorbed hydrogen inside the zirconium matrix is in solid solution at 20 C 
(<100 ppm), the El 10 has very high ductility at 20 C. Therefore the increase of ductility margin does not 
occur in the range 20-135 C; 

• the critical value of hydrogen content associated with the El 10 zero ductility threshold is increased from 
400 ppm at 20 C up to 900 ppm at 135 C. This effect is most likely associated with the increase of hy¬ 
dride ductility; 

• the embrittled sample with hydrogen content higher than 1000 ppm is insensitive to the temperature in 
this range. 

Several ring compression and ring tensile tests performed at the temperature 200-300 C have shown that the 
cladding ductility is sharply increased in the temperature range 20-200 C with any hydrogen content 
(1500 ppm is the maximum value). The temperature increase up to 300 C does not change the ductility mar¬ 
gin in general. 

Therefore, the results of these investigations allowed to conclude that: 

• the Zry-4 embrittlement is caused by the oxygen absorption in the prior (3-phase and by the reduction in 
prior P-phase thickness; 

• the E110 embrittlement is a function of two processes: oxygen embrittlement and hydrogen embrittle¬ 
ment accompanied by the reduction in the prior p-phase thickness. 

To estimate the representativity of obtained results for other oxidation temperatures, the additional oxidation 
tests were performed in the range 800-1200 C. This stage of investigations has shown that: 

• the most pronounced effects of the breakaway oxidation were observed at the temperatures 950-1000 C 
(see Fig. 5.8). The zero ductility threshold in this temperature range is a little lower than that at 1100 C 
(7.5% ECR at 1000 C); 

• in spite of the presence of breakaway effects at the temperature 800 C, the zero ductility threshold is 
increased at that temperature to more than 12% ECR. This conclusion is confirmed by the low hydrogen 
content in the oxidized cladding. The critical hydrogen concentration (about 400 ppm) is achieved at 


5.7 


about 12% ECR. A detailed analysis of appropriate physical processes is a task for the future research 
but preliminary it should be noted that this improved behavior of the El 10 cladding can be associated 
with the two phase compositions of the zirconium matrix (a and P-phases, in this case, the concentration 
of P-phase is relatively low at this temperature), a very low thickness of a-Zr(O) at this temperature, and 
some other phenomena considered in the report; 

• as for the temperature 1200 C, the zero ductility threshold of the El 10 cladding is most likely not better 
than that at 1100 C in spite of the tendency towards the decrease of hydrogen absorption rate. 


Fig. 5.8. Demonstration of the El 10 breakaway oxidation effects at 950 C 

To determine the sensitivity of the embrittlement behavior of niobium-bearing claddings to such alloying 
components as Sn and Fe, special investigations were performed with the E635 cladding (Zr-l%Nb-l .2%Sn- 
0.35%Fe). The analysis of obtained results showed that the visual appearance of E635 oxidized claddings 
was better than that of El 10 claddings but in spite of this fact, the zero ductility threshold for these cladding 
materials was practically the same. 

The last line of experimental investigations performed with the as-received El 10 cladding tube was con¬ 
nected with the determination of the sensitivity of oxidation and mechanical behavior to the initial oxygen 
concentration in niobium-bearing alloys. The motivation to perform these studied was associated with the 
absence of a precise position on this issue in the current scientific publications and with the fact that the ini¬ 
tial oxygen concentration in the tested El 10 cladding was very low (~0.05%) in comparison with other al¬ 
loys (including other niobium-bearing alloys). To develop the comparative data base, E110K as-received 
cladding tubes were tested. The initial oxygen concentration in this modification of the El 10 alloy was about 
0.11% by weight. The obtained test results allowed to establish that: 

• the increase of oxygen concentration in the El 10 alloy does not lead to suppression of the breakaway 
oxidation effect; 

• the zero ductility threshold of the El 10K cladding is not higher than that of the standard El 10 cladding. 

5.1.4. The embrittlement behavior of Zr-l%Nb (El 10) cladding 
as a function of irradiation effects 

Taking into account that the validation of high burnup fuel behavior under LOCA conditions is one of the 
urgent research problems, the preliminary stage of this type of investigations was performed in the frame of 
this work. 

The irradiated El 10 claddings refabricated from VVER high burnup fuel rods with burnup of about 
50 MW d/kg U had the following characteristics before the oxidation tests: 

• Zr0 2 thickness on the outer cladding surface was 5 pm; 

• Zr0 2 thickness on the inner cladding surface was 0 pm; 

• hydrogen content in the cladding material was 47 ppm. 


• E110 as-received tube 

• F/F combination of heating and cooling rate 

• ECR=11.2% 


• Microstructure of oxidized sample => 

• Appearance of oxidized sample li 




5.8 










The analysis of the appearance and metallographic examinations of the irradiated cladding oxidized in the 
temperature range 1000-1200 C at the FF combination of heating and cooling rates allowed to reveal the 
following (see Fig. 5.9): 


the outer surface of oxidized irradiated claddings are covered with the black uniform oxide without the 
spallation effects up to 16% ECR based on the results of visual examinations: 

as for the inner surface, the tendency towards oxide spallation was noted at the 7.7% ECR and higher. 


Before the tests 


ECR=0.5% 


isspsew; 






1100C 


ECR-7.7 % (Polished) 


1 * • '•* . &'• %**•«*• »A'» V* . 


Outer surface 


After the oxidation at 1200 C 


ECR=16% 



Inner surface 


Fig. 5.9. The appearance and microstructure of the El 10 irradiated claddings 

as a function of the ECR 


The obtained data are in a good agreement with other Russian investigations performed in the MIR research 
reactor under LOCA conditions. The following general observations may be made on the basis of the whole 
scope of experimental data: 

• the irradiation inhibits the breakaway oxidation tendency on the outer surface of the El 10 cladding; 

• the tendency towards oxide spallation is supplemented with the tendency towards an increase in 
oxidation rate (the oxide thickness increase) on the cladding inner surface. In accordance with these data, 
it may be assumed that the contamination of the cladding inner surface by fission products is responsible 
for these effects. 

The consideration of the data base characterizing the residual ductility microhardness and hydrogen content 
in the oxidized irradiated cladding as a function of the ECR has shown that: 

• in accordance with the postulated relationship between the oxygen concentration in the prior (3-phase, 
microhardness, and the oxygen induced zero ductility threshold, this threshold corresponds to 8.3% 
ECR: 

• the combination of the oxygen induced and hydrogen induced embrittlement of the El 10 irradiated 
cladding leads to the reduction of the zero ductility threshold down to 6.5% ECR at the F/F combination 
of heating and cooling rates. 

5.1.5. The oxidation kinetics and embrittlement behavior of the El 10 cladding 
according to results of previous investigations in different laboratories 

Earlier oxidation and ring compression tests with the El 10 unirradiated cladding were performed in the 
following institutes: 

• VNIINM. Russia [1]; 

• KFKI. Hungary [2, 3]; 

• NFI, Czech republic [4, 5]; 

• NC in Rossendorf. Germany [6, 7]. 


5.9 










It should be noted that the comparative analysis of VNIINM test results was performed on the basis of data 
obtained at the end of 1980s of the past century. The recent VNIINM tests were not used because open 
VNIINM publications devoted to these issues did not contain the information in detail concerning test modes 
and test parameters. 

The comparison of the El 10 oxidation kinetics estimated according to the results of this work with oxidation 
kinetics obtained earlier has shown that the discrepancy between the RRC KI/RIAR, VNIINM, NFI, NC in 
Rossendorf results is low in the studied range of weight gains and temperatures (see Fig. 5.10). The KFKI 
data overestimates noticeably the El 10 oxidation kinetics. 



Fig. 5.10 The comparison of the El 10 oxidation kinetics (1073-1473 K) in accordance 

with the data of different investigations 

The comparison of the El 10 and Zry-4 oxidation kinetics presented in Fig. 5.11 for the as-measured data and 
for the as-calculated data (with the use of Zry-4 conservative kinetics based on the Baker-Just correlation 
and E110 conservative kinetics developed in the VNIINM) allows to note the following: 

• practically the same conservative kinetics is used for the safety analysis of the El 10 and Zry-4 claddings; 

• the as-measured El 10 oxidation kinetics is noticeably less than that for the Zry-4 cladding. 

The difference in the postulated safety criteria characterizes the El 10 and Zry-4 fragmentation thresholds 
(18% and 17% respectively) and the difference in the El 10 and Zry-4 oxidation kinetics leads to the 
following estimations of the as-measured oxidation corresponding to safety criteria at 1100 C: 

• 11.5% ECR for El 10; 

• 13.5% ECR for Zry-4. 

The preliminary data characterizing the oxidation kinetics of the El 10 irradiated cladding showed that the 
oxidation rate of irradiated claddings was somewhat higher than that of unirradiated claddings at all oxida¬ 
tion temperatures (1000-1200 C). But taking into account the limited number of these tests, the investiga¬ 
tions in this line should be continued in the future. 

The analysis of comparative data characterizing the embrittlement behavior of the El 10 unirradiated 
claddings allowed to note that all investigators revealed the tendency towards the breakaway oxidation of the 
E110 cladding accompanied by the hydrogen absorption and a sharp reduction in the residual ductility 
margin after the initiation of oxide spallation. 

The consideration of the whole scope of the El 10 test data obtained in the first part of this research led to the 
decision concerning the development of the program second stage to be devoted to the analysis of reasons for 
the El 10 specific behavior in comparison with other niobium-bearing alloys. 

5.10 
























Fig. 5.11. The comparison of the El 10 (on left) and Zry-4 (on right) conservative and as-measured 

kinetics at 1100 C 


5.2. Major findings of the program second part 


5.2.1. The concept of special investigations 

The following general questions were formulated on the basis of the revealed general difference in the 
oxidation and embrittlement behavior of Zr-l%Nb (El 10) and Zry-4 claddings: 

Is the earlier breakaway initiation and the embrittlement caused by combined effects of oxygen and 
hydrogen uptake typical of the whole family of niobium-bearing alloys or does this phenomenon 
characterize the El 10 alloy only? 

The beginning of investigations to provide the answer to this question was devoted to the comparison of Zr- 
l%Nb (El 10) embrittlement behavior with the behavior of other Zr-l%Nb claddings manufactured from the 
M5 alloy [12]. Published data with experimental results on the M5 cladding available by that time were used 
for this goal [13, 14], The analysis of these M5 test results has shown that: 

• the zero ductility threshold of the M5 cladding is higher than that of the El 10 cladding; 

• the breakaway oxidation is not observed in the investigated range of parameters; 

• the M5 embrittlement is not accompanied by hydrogen uptake. 

The analysis of possible reasons for the revealed difference between El 10 and Zry-4 or M5 cladding behav¬ 
ior allowed to select the following phenomena for special studies: 

• surface effects; 

• bulk effects associated with the chemical composition of impurities in the cladding material; 

• bulk effects associated with the cladding microstructure. 

This part of the research program was closely coordinated with another experimental program being con¬ 
ducted at ANL at approximately the same time [15, 16], The oxidation and mechanical tests were performed 
with M5, Zirlo, El 10, and Zry-4 claddings in that program. The results of both programs allowed an exten¬ 
sion of the comparative data base for observations and conclusions. 

5.2.2. Surface effect studies 

These studies were based on the following considerations: 


5.11 












































• the corrosion resistance is a function of the cladding surface finish because the corrosion behavior 
depends on the surface chemistry (contaminations) and surface roughness; 

• the sensitivity of the El 10 cladding material to these phenomena should be estimated. 

The urgency of this issue was determined by the fact that as-received El 10 tubes used in the first part of the 
program were not subjected to surface finishing. The current El 10 surface finishing procedure consisted of 
the final etching and anodizing of the cladding outer surface. To determine the sensitivity of the cladding 
embrittlement behavior to this procedure, the oxidation and mechanical tests with the etched and anodized 
El 10A cladding were performed. The results of tests have shown that this surface finishing procedure does 
not allow to eliminate the earlier breakaway oxidation and to improve the embrittlement behavior of the 
E110 cladding. 

Besides, one of the advanced methods for the surface finishing with the use of the outer surface grinding and 
inner surface jet etching was investigated also. But the conclusion was the same. Moreover, the tendency 
towards an increase in oxide thickness was revealed on the cladding inner etched surface. These results were 
in a good agreement with the ANL test data obtained with the El 10 as-received tubes subjected to the special 
etching. The ANL tests showed that any etching led to the degradation of the cladding embrittlement behav¬ 
ior. 


To eliminate the chemical contamination of the cladding surface caused by etching, special polished El 10 
samples were used for the next stage of the surface effect studies. The results of tests with this type of the 
cladding samples allowed to reveal the following (see Fig. 5.12): 

• the visual indicators of the breakaway oxidation practically disappeared from polished parts of oxidized 
samples. The most pronounced effect was noted at 1000 C; 

• the polished and unpolished parts of the oxidized cladding were characterized with quite a different 
hydrogen uptake; 




the residual ductility increased very significantly on the polished part of the oxidized cladding. 


The same results have been obtained in similar investigations performed at ANL [15, 16]. 



Unpolished 


Polished 


Hydrogen content: 
173 ppm 


Hydrogen content: 
44 ppm 


Fig. 5.12. The appearance of the El 10 polished and unpolished parts after the oxidation at 1000 C 


Thus, the surface polishing allows to improve the oxidation and embrittlement behavior of the niobium¬ 
bearing cladding of the El 10 type. 


5.2.3. Bulk chemistry studies 

Several investigations performed during last years have demonstrated that the oxidation behavior of 
niobium-bearing alloys is very sensitive not only to alloying components but also to the impurity 
composition. The analysis of the El 10 problems in this context has shown that different methods are used to 
produce zirconium alloys for PWR and VVER claddings. In accordance with these methods, the sponge Zr is 
used to produce such alloys as Zry-4, M5, Zirlo and the mixture of iodide and electrolytic Zr is used for the 
fabrication of the El 10 alloy. 

It is obvious that the difference in the alloy production leads to the difference in the impurity compositions. 
This fact was accepted for the basis while developing a special subprogram devoted to the oxidation and ring 
compression tests with modified types of the El 10 claddings (El 10 q). 

These modified types of the El 10 claddings are the pilot samples produced by the Russian industry in accor¬ 
dance with the program of the sponge Zr introduction in the production of VVER claddings. The oxidation 
tests performed at 1100 C with different variants of the El 10 sponge type claddings have demonstrated prac- 


5.12 








tically the same result. Macroscopic effects of the breakaway oxidation mode disappeared in the as-measured 
ECR range 10.5-18% (see Fig. 5.13). 


Iodide/ electrolytic 
E110 

ECR-16% 


Sponge E110 

ECR=18% 



Fig. 5.13. The comparison of the iodide/electrolytic 
and sponge El 10 cladding oxidation behavior at 1100 C 


For oxidation at 1100 C, ring compression tests with oxidized claddings confirmed that the zero ductility 
threshold of the sponge El 10 increased up to 19% ECR (as-measured) due to the low hydrogen uptake. The 
preliminary comparison of the embrittlement behavior of the sponge El 10 and other alloys oxidized at 
1100 C with the ANL published data on Zry-4, Zirlo. M5 test results [16] showed approximately the same 
zero ductility threshold. 

Nevertheless, some difference was revealed in the hydrogen uptake for these claddings at 1100 C. Thus, 
Zry-4, Zirlo and M5 were characterized by the hydrogen content of 17-22 ppm at 19.1-21.1% ECR [16]. 
The E110 cladding kept the same hydrogen content (17 ppm) up to 16.7% ECR. But at 18% ECR. the hy¬ 
drogen content increased up to 102 ppm. To determine the sensitivity of the sponge El 10 to the oxidation 
temperature, special investigations were performed in the range 900-1200 C. These tests allowed to reveal 
several new phenomena. 

For oxidation at 900 - 1000 C, major new phenomena were revealed and are associated with a sharp reduc¬ 
tion of the sponge El 10 oxidation rate in comparison with that of the iodide/electrolytic El 10 (see Fig. 5.14), 
although at 1100 - 1200 C the oxidation kinetics of the sponge El 10 and iodide/electrolytic El 10 were simi¬ 
lar with some tendency towards an increase of the oxidation rate in the sponge El 10 claddings. The compari¬ 
son of the sponge El 10 with the other sponge type of Zr-l%Nb cladding (M5 alloy) performed with the use 
of French data [17] showed that oxidation kinetics of these alloys were the same at both temperatures. 



Fig. 5.14. The comparison of the Zry-4, sponge El 10, iodide/electrolytic El 10, M5 

oxidation kinetics at 1000 C 

The comparison of obtained data with the Zry-4 oxidation kinetics is presented as a function of temperature 
in Fig. 5.15. 


5.13 



























10 


c 

o 

<N 

ao 

£ 

<u 
*—< 
C3 

1— 

e 

o 

c5 

JO 

x 

O 



0.01 1 


0.001 


0.0001 


0.001 


0.0006 0.0007 0.0008 0.0009 

1/T (K 1 ) 

Fig. 5.15. The comparison of the Zry-4, sponge El 10, iodide/electrolytic El 10 
oxidation rates in the temperature range of 900 - 1200 C 


The data base obtained at 900 C and 1000 C to characterize the embrittlement behavior of the sponge El 10 is 
not sufficient for the determination of the zero ductility threshold. Nevertheless, the analysis of test results 
allowed to make the following important conclusions: 

• the ductility reduction as a function of ECR is caused by the oxygen induced mechanism at the low 
hydrogen uptake; 

• the critical ECR associated with the zero ductility threshold at 900 C and 1000 C will be significantly 
lower than that at 1100 C, although the critical time will be significantly higher; 

• a possible explanation of this phenomenon can be associated with the formation at 1000 C of a thicker 
a-Zr(O) layer in comparison with the a-Zr(O) layer formed at 1100 C or in comparison with the a-Zr(O) 
layer in the iodide/electrolytic El 10 formed at 1000 C. This effect leads to the reduction of the effective 
thickness in the prior P-phase layer and to the increase of oxygen content in the metallic matrix; 

• besides, the first indications of the breakaway oxidations (white spots on the cladding) appear at the 
ECR>8.5%, but this corresponds to a very long time. 

It should be noted in the context of the revealed features that similar tendencies were noted in other Zr-l%Nb 
cladding alloys. Thus, a sharp reduction of M5 ductility was observed in the ECR range 11-12% (as- 
measured) in the ANL test data [15, 16]. Moreover, in accordance with results of French investigations, the 
breakaway phenomena accompanied by high hydrogen and nitrogen pickup were noted at 1000 C under the 
following conditions: the oxidation time was much higher than 1800 s but less than 135000 s [18]. It is of 
interest that the ANL data showed that Zirlo cladding did not demonstrate the tendency towards the earlier 
ductility reduction at 1000 C and it did not demonstrate a tendency of a sharp reduction in oxidation rate 
[15]. 

The next important data characterizing the sponge El 10 behavior were obtained at the oxidation temperature 
1200 C. The overview of appropriate results is as follows: 

• in spite of the fact that visual indications of the breakaway oxidation were not observed up to the 23.3% 
ECR, a significant hydrogen uptake was revealed already at 8% ECR; 

• the zero ductility threshold of the sponge El 10 did not exceed the 8% ECR at 1200 C. 

The comparison of these data with the ANL data [15, 16] characterizing the Zry-4, M5, Zirlo embrittlement 
behavior after the oxidation at 1200 C showed that: 


5.14 
















• all tested alloys showed a tendency towards the reduction of the zero ductility threshold at this 
temperature. The Zry-4, Zirlo, M5 ductility reduction down to 5% (the residual ductility or offset hoop 
strain) occurred in the measured ECR range 8.2-10% ECR; 

• but the Zry-4. Zirlo, M5 embrittlement was not accompanied by high hydrogen uptake. The hydrogen 
content in these claddings remained at the level ranging from 17-19 ppm up to 18.8-22.3% ECR. 

The consideration of iodide/electrolytic and sponge El 10 comparative data allows to assume that revealed 
general differences in the behavior of these cladding materials are a function of differences in the micro¬ 
chemical composition of these two modifications of the El 10 alloy. To clarify the background of the prob¬ 
lem. the results of previous investigations devoted to the relationship between the corrosion resistance of 
zirconium-based claddings and microchemical composition of the cladding alloy were reassessed. The major 
outcomes of this work w ere the following: 

• the stabilization of zirconium dioxide tetragonal form led to the improvement of the cladding corrosion 
resistance; 

• in this context, all impurities can be subdivided into the beneficial and deleterious impurities; 

• in accordance with the physical theories and empirical data, the beneficial and deleterious impurities 
consisted of the following elements: 

- beneficial impurities: Fe, Cr, Ca, Mg, Y... 

- deleterious impurities: C, N, F, Cl, Si, Ti, Ta. V, Mn, Pt, Cu... 

• there are contradictory points of view r regarding such elements as Al, Ni, Mo. As for oxygen, the major¬ 
ity of investigators consider that this element is neutral with respect to corrosion resistance; 

• the corrosion behavior w r as very sensitive to the concentration of such alloying elements as Nb and Sn. 
Moreover, each type of alloy had the optimal concentration of the alloying element at which the best 
corrosion resistance was provided: 

• the main potential differences between the microchemical compositions of the iodide/electrolytic and 
sponge E110 alloy may be associated with the following: 

- the method of the electrolytic El 10 production led to the risk of the alloy enrichment with such 
deleterious impurity as fluoride. Besides, the electrolytic El 10 had a very high Hf content, the role of 
which was not quite understood with respect to the cladding corrosion behavior. The iodide Zr 
component of the alloy may be considered as neutral because this component had a very low content of 
both beneficial and deleterious impurities; 

- the method of the sponge El 10 production facilitated the alloy enrichment with such beneficial 
impurities as Ca, Mg, Fe, Y. 

Taking into account the results of this analysis, chemical compositions of iodide/electrolytic and sponge 
E110 alloys were compared. The comparison showed that reasonable differences in the impurity contents 
were revealed for two elements only: 

• Fe: 86 ppm in the iodide/electrolytic El 10 and 120-140 ppm in the sponge El 10; 

• Hf: 350 ppm in the iodide/electrolytic El 10 and 90-420 ppm in the different modifications of the sponge 
E110. 

It should be noted that it was impossible to compare directly the contents of many impurities of a low content 
in the alloy due to very low concentrations: the content was less than 10 ppm, 30 ppm, etc. 

To determine the sensitivity of the El 10 oxidation behavior, special tests were performed using the 
iodide/electrolytic El 10 cladding with a low hafnium content (El 10 towH f, 90 ppm). These tests showed that 
the corrosion resistance was a function ot halnium content in the studied range of Hf variation (350 ppm in 
the iodide/electrolytic El 10 and 90 ppm in the El 10| OVV Hf)- 

The El 10| OWH f cladding was characterized by the increase of the zero ductility margin from 8.3% ECR 
(standard El 10) up to 12% ECR due to the significant delay in the breakaway initiation as a function of 
ECR. However, the fact cannot be ruled out that in this case the process of the El 10 alloy purification from 

5.15 


hafnium might be accompanied by the change in the content of any other impurity because some 
modifications of the sponge El 10 tested in the frame of this work constituted the mixture of sponge Zr, 
iodide Zr and recycled scrap with a high hafnium content but the corrosion resistance of these modifications 
was very high. 

The next position of special investigations was associated with the determination of the sensitivity of the 
oxidation behavior of niobium-bearing alloys to the iron content: 

• the oxidation and mechanical tests of the E635 cladding with a very high iron content in the cladding 
material (0.34-0.4% by weight) showed that the oxidation behavior and embrittlement threshold of 
iodide/electrolytic E635 was somewhat better than that for the iodide/electrolytic El 10 but the sponge 
variant of E635 had practically the same embrittlement characteristics as iodide/electrolytic E635. And 
both versions of the E635 alloy had lower zero ductility thresholds (at 1100 C) than the sponge El 10; 

• the analysis of special investigations performed by the VNIINM [19] with the variation of iron content in 
the range 80-1400 ppm and the variation of such elements as O, C, Hf, Cr showed that: 

- in spite of the fact that the oxidation and mechanical response of seven types of the oxidized cladding 
(10% ECR) differentiated from the best (lustrous black oxide, low hydrogen content, reasonable margin 
of residual ductility) to the worst (breakaway effects, oxide spallation, high hydrogen content, low 
residual ductility), a direct association between the appropriate response and combinations of the 
chemical composition limited with such elements as Nb, Fe, C, Hf, Cr, O was not managed to be 
revealed. This fact undoubtedly indicated that the influence of other varied elements or some other 
parameters was not taken into account; 

- nevertheless, it should be noted that the best corrosion behavior of the tested cladding was obtained with 
an iron content of about 130-450 ppm and low hafnium content (~ 100 ppm) and a low carbon content. 
Very low (80 ppm) and very high (1400 ppm) iron contents were associated with the intermediate 
corrosion behavior but were also associated with very high carbon content (up to 200 ppm) in several 
samples. 

Finally, the following general recommendations concerning the relationship between the microchemical 
composition and embrittlement behavior of niobium-bearing alloys can be given: 

• the characterization of types of niobium-bearing alloys based on the current list of alloying elements 
should be reassessed; 

• the status of some impurity elements must be changed and the definition of minor alloying elements 
should be added to the characterization of the cladding alloy; 

• special investigations to determine the deleterious, beneficial and neutral impurities in the niobium¬ 
bearing alloys should be performed additionally. 

5.2.4. Bulk microstructure studies 

The last position of this stage of the research program was connected with the comparison of the 
microstructure parameters of different types of El 10 cladding. To develop the data base for the comparison, 
TEM examinations of iodide/electrolytic El 10 and sponge El 10 cladding materials were performed. 

In accordance with the results of analytical studies, the following comparative data were measured: 

• phase conditions and phase composition; 

• the grain size of zirconium in the cladding matrix; 

• the parameters of the secondary phase precipitates being of importance for the corrosion resistance: the 
chemical composition, size, density, uniformity of distribution. 

The choice of these comparative data was based on the results of inside and outside Russian investigations 
devoted the determination of the relationship between the corrosion behavior and cladding microstructure. 

The results of TEM investigations showed that: 


5.16 


• iodidc/electrolytic El 10 and sponge El 10 had a completely recrystallized microstructure (a+P-Nb); 

• the average size of a-Zr grain in the matrix was similar for both types of El 10 claddings (2.8 pm for 
iodide/electrolytic El 10 and 3.2 pm for the sponge El 10); 

• the globular type of P-Nb precipitates uniformly distributed in the a-Zr matrix characterized both El 10 
modifications; 

• the average size of the secondary P-Nb precipitates was practically the same in the iodide/electrolytic and 
sponge E110 (41-60 pm); 

• the iodide/electrolytic El 10 did not contain the intermetallic precipitates; the sponge El 10 had the 
intermetallic precipitates of the Zr(Nb,Fe )2 type, and the size of these precipitates was about 180 pm. 

The comparative analysis of the iodide/electrolytic El 10 and sponge El 10 microstructures with the French 

published data on the M5 microstructure allowed to conclude the following: 

• iodide/electrolytic El 10 (El 10i ow Hf), sponge El 10 and M5 had practically the same parameters for the 
microstructure; 

• the only distinction between the microstructure of iodide/electrolytic El 10 and sponge types of Zr-l%Nb 
cladding (sponge El 10, M5) was that the sponge Zr-l%Nb types of alloy had the iron-based precipitates 
in addition to the P-Nb precipitates. 

5.2.5. Final Remarks 

Taking into account the results of investigations devoted to the microstructure, microchemical and surface 

effects, the following general conclusions can be made: 

• the current type of the El 10 (Zr-l%Nb) cladding (standard iodide/electrolytic) has quite an optimal mi¬ 
crostructure. The specific oxidation behavior of this cladding at high temperatures is not a function of the 
cladding microstructure; 

• the performed research allowed to establish that the oxidation behavior and ductility margin of the El 10 
oxidized cladding are very sensitive to the cladding microchemical composition and surface finishing; 

• the use of the sponge type of zirconium for the fabrication of cladding tubes provides a significant reduc¬ 
tion of the cladding oxidation rate, especially in the temperature range of 900 - 1000 C, and an increase 
in the zero ductility threshold; 

• additional improvement of the oxidation behavior and a significant increase of the residual ductility mar¬ 
gin in the El 10 oxidized cladding can be achieved by polishing of the outer and inner cladding surfaces. 


5.17 


References for Section 5 


[1] Bibilashvili Yu.K., Sokolov N.B., Dranenko V.V., Kulikova K.V., Izrailevskiy L.B., Levin A.Ya., Moro¬ 

zov A.M. "Influence of Accident Conditions due to Loss of Tightness by Primary Circuit on Fuel Clad¬ 
dings", Proc. of the Ninth Int. Symp. on Zirconium in the Nuclear Industry \ Kobe, Japan, November 5-8, 
1990 (ASTM STP-1132, 1991). 

[2] Hozer Z., Griger A., Matius L., Vasaros L., Horvath M. "Effect of Hydrogen Content on the Embrittle¬ 

ment of ZR Alloys", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Transient 
and LOCA Conditions ", Halden, Norway, September 10-14, 2001. 

[3] Hozer Z., Matus L., Horvath M., Vasaros L., Griger A., Maroti L. “Ring Compression Tests with Oxi¬ 

dized and Hydrided Zr-l%Nb and Zirkaloy^t Cladding”, Hungarian Academy of Sciences, CRIP, Bu¬ 
dapest, Report KFKI-2002-01/G. 

[4] Vrtilkova V., Valach M., Molin L. "Oxiding and Hydrating Properties of Zr-l%Nb Cladding Material in 

Comparison with Zircaloys", Proc. of IAEA Technical Committee Meeting on "Influence of Water 
Chemistry> on Fuel Cladding Behavior ", Rez (Czech Republic), October 4-8, 1993. 

[5] Vrtilkova V., Novotny L., Doucha R., Vesely J. “An Approach to the Alternative LOCA Embrittlement 

Criterion”, Proc. of SEGFSM Topical Meeting on LOCA Fuel Issues, Argonne National Laboratory, 
May 2004 (NEA/CSNI/R(2004)19). 

[6] Bohmert J. "Embrittlement of Zr-l%Nb at Room Temperature after High-Temperature Oxidation in 

Steam Atmosphere", Journal, Kemtechnik 57 Nol, 1992. 

[7] Bohmert J., Dietrich M., Linek J. "Comparative Studies on High-Temperature Corrosion of Zr-l%Nb and 

Zircaloy-4", Nuclear Engineering and Design, 147 Nol, 1993. 

[8] Bibilashvili Yu.K., Sokolov N.B., Salatov A.V., Andreyeva-Andriyevskaya L.N., Nechayeva O.A., 
Vlasov F.Yu. "RAPTA-5 Code: Modelling of Behaviour of Fuel Elements of VVER Type in Design 
Accidents. Verification Calculations", Proc. of IAEA Technical Committee Meeting on "Behavior of 
LWR Core Materials under Accident Conditions", Dimitrovgrad, Russia, on 9-13 October 1995. IAEA- 
TECDOC-921, Vienna, 1996. 

[9] Gyori Cs. et.al. “Extension of Transuranus code applicability with niobium containing models 
(EXTRA)”. Proc. of FISA-2003 Conference EU Research in Reactor Safetv, EC Luxembourg, Novem¬ 
ber 2003. 

[10] Baker L., Just L.C. “Studies of Metal Water Reactions at High Temperatures”, Technical report of 
ANL-6548, 1962. 

[11] Cathcart J.V. and Pawel R.E. “Zirconium Metal-Water Oxidation Kinetics: IV Reaction Rate Studies”, 
ORNL/NUREG-17, 1977. 

[12] Mardon J.P., Gamer G., Beslu P., Charquet D., Senevat J. “Update on the Development of Advanced 
Zirconium Alloys for PWR Fuel Rod Claddings”. Proceedings of the International Topical Meeting on 
Light Water Reactor Fuel Performance, Portland, Oregon, March 2-6, 1997. 

[13] Brachet J., Pelchat J., Hamon D., Maury R., Jaques P., Mardon J. "Mechanical Behavior at Room Tem¬ 
perature and Metallurgical Study of Low-Tin Zry-4 and M5™ (Zr-NbO) Alloys after Oxidation at 
1100°C and Quenching", Proc. of IAEA Technical Committee Meeting on "Fuel Behavior under Tran¬ 
sient and LOCA Conditions", Halden, Norway, September 10-14, 2001. 

[14] Mardon J., Frichet A., Bourhis A. "Behavior of M5™ Alloy under Normal and Accident Conditions", 
Proc. of Top Fuel-2001 Meet., Stockholm, May 27-30, 2001. 

[15] Billone M.C., Yan Y. and Burtseva T. “Post-Quench Ductility of Zircaloy, El 10, ZIRLO and M5”, 
Proc. of SEGFSM Topical Meeting on LOCA Fuel Issues, Argonne National Laboratory, May 2004 
(NEA/CSNI/R(2004)19). 

[16] Billone M.C., Yan Y. and Burtseva T. “Post-Quench Ductility of Advance Alloy Cladding”, 2004 Nu¬ 
clear Safety Research Conference, Washington DC, October 2004. 


5.18 


[17] Mardon J.P., Waeckel N. “Behavior of M5™ Alloy under LOCA Conditions'’, Proc. of Top Fuel-2003 
meetings “Nuclear Fuel for Today and Tomorrow Experience and Outlook ”, March 16-19, 2003, 
Wurzburg, Germany. 

[18] Waeckel N., Mardon J.P. “Recent data on M5™Alloy under LOCA Conditions”, Proceedings of the 
2003 Nuclear Safety Research Conference , Washington DC, October 20-22, 2003, NUREG/CP-0185, 
2004. 

[19] Nikulina A.V., Andreeva-Andrievskaya L.N., Shishov V.N., Pimenov Yu.V. “Influence of Chemical 
Composition of Nb Containing Zr Alloy Cladding Tubes on Embrittlement under Conditions Simulating 
Design Basis LOCA'’, Fourteenth International Symposium on Zirconium in the Nuclear Industry \ 
ASTM STP. 


5.19 






































APPENDIX A 

Description of Test Apparatus and Test Procedures 


A-1 






A-l. Description of oxidation test apparatus 


The scheme of oxidation apparatus is presented in Fig. 3.2 (see section 3). The oxidation facility consists of the following 
basic elements: 

• electric furnace; 

• device for the movement of El 10 sample in the several given positions (cold position, hot position, quench position); 

• steam generator; 

• system to supply the argon to the oxidation facility; 

• measurement apparatus (measurement of temperature inside and outside of El 10 sample, temperature of water in the 
steam generator); 

• tank with the water in the low part of facility. 

The electric furnace provided the radiant heating and cooling of cladding sample with the different temperature rates (see 
the more detail information in Appendix A-2). The steam generator provided the generation of water steam with the fol¬ 
lowing parameters: 

• temperature: 150 C; 

• mass flow rate: 0.01-0.04 g/s (mass flow rate is a function of electric power); 

• atmospheric pressure. 

The system to supply the argon to the oxidation facility was used at the beginning of each test mode. This system devoted 
the cleaning of gas atmosphere inside the test facility from air. The temperature limit was used to change the type of cool¬ 
ant. The Ar flow was stopped at the temperature about of 300 C and after that the heating was continued with the water 
steam. 

The important details of temperature measurements are described in Appendix A-4. The tank with water in the low part of 
facility devoted the opportunity to perform cooling of cladding sample under quench conditions. 


A-2. Description of oxidation procedures 


The five combinations of heating and cooling rates were used for this experimental program (see Fig. 3.3 in section 3): 

1. Slow heating and slow cooling (S/S). 

2. Slow heating and fast cooling (S/F). 

3. Fast heating and slow cooling (F/S). 

4. Fast heating and fast cooling (F/F). 

5. Fast heating and quench cooling (F/Q). 

The characterization of major types of test modes is described below. 

Slow heating (Fig. A-l): 

• the cladding sample was installed in the hot position (see Fig. 3.2 of section 3); 

• the air atmosphere in the electric furnace was replaced on the argon atmosphere; 

• the electric furnace was switched on and the cladding sample was heated to 150 C with the temperature rate 0.5 C/s; 

• the argon atmosphere was replaced on the water steam atmosphere and the heating of cladding sample was continued 

with the rate 0.5 C/s to the hold temperature. 


A-2 



; 


o 


» w 

J— 

O 

c. 

g 

_CJ 

c. 

g 

C3 

C/5 



Time 


Fig. A-l. Oxidation mode with the “slow” heating: sequence diagram 


Fast heating (Fig. A-2): 

• the cladding sample was installed in the cold position (see Fig. 3.2 of section 3); 

• the air atmosphere in the electric furnace was replaced on the argon atmosphere; 


• the electric furnace was switched on and the gas medium inside the electric furnace (in the hot position) was heated to 
150 C after that the argon atmosphere was replaced on the water steam and the heating of the furnace was continued 
during 600 seconds to the hold temperature (in accordance with the thermocouple #2 data (see Fig. 3.2 of section 3)), 
the temperature of cladding sample, which was located in the cold position was not above 150 C at the end of this 
stage; 

• the cladding sample was moved in the hot position the such a way to provide the heating rate 25 C/s in the tempera¬ 
ture range 150-800 C. 





c. 

£ 

C/5 



Time 


Fig. A-2. Oxidation modes with “slow”, “fast”, “quench” cooling: sequence diagram 


A-3 



























































Different types of cooling (Fig. A-3): 

• As the slow cooling the cladding sample was cooled in the hot position with the temperature rate 0.5 C/s to 150 C, 
after that the steam atmosphere was replaced on the argon atmosphere and the cooling of cladding sample was contin¬ 
ued with the same cooling rate. 

• At the fast cooling the cladding sample was moved from the hot position to the cold position the such a way that the 
cooling rate at initial phase of this process was 25 C/s. The replacement of steam coolant to the argon coolant was 
made at the temperature of cladding sample 150 C. 

• At the quench cooling the cladding sample was moved from the hot position to the quench position (to the cold water) 
with cooling rates about of 200 C/s. 


<D 


C3 

s— 

o 

Cl 

E 

a 
.— 1 

jj 

c. 

£ 

a 

on 



A BCD 


Time 


Fig. A-3. Oxidation mode with the “’fast” heating: sequence diagram 


A-3. The characterization of cladding samples 


Table A-l. The list of tested cladding materials 


Abbreviation 

Alloying composition 

Additional comments 

E110 

Zr-l%Nb 

as-received El 10 cladding tube manufactured from the iodide Zr, 
electrolytic Zr, and recycled scrap (the oxygen concentration 
~0.04 % by weight) 

E110A 

Zr-1 %Nb 

as-received El 10 cladding after the etching and anodizing of as- 
received E110 cladding tube 

E110K 

Zr-l%Nb 

as-received El 10 cladding tube manufactured from the iodide Zr, 
electrolytic Zr, and recycled scrap with the high content of oxygen 
concentration (0.11 % by weight) 

Ell 0i ow Hf 

Zr-1 %Nb 

as-received El 10 cladding tube manufactured from the iodide Zr, 
electrolytic Zr, and recycled scrap with the low content of Hf 
(90 ppm). The standard content of Hf is about of 350 ppm 

E635 

Zr-1 %Nb-1.2%Sn- 
0.35%Fe 

as-received E635 cladding tube manufactured from the iodide Zr, 
electrolytic Zr, and recycled scrap 

Zry-4 

Zr-1.4%Sn 

the cladding tube manufactured by Framatom ANP GmbH on 
January 1989 


A-4 








































Abbreviation 

Alloying composition 

Additional comments 

El 10 irr 

Zr-1 %Nb 

the irradiated El 10 cladding refabricated from the commercial 
VVER-1000 fuel elements irradiated to 50-53 MWd/kg U 

El lOo(fr) 

Zr-l%Nb 

as-received El 10 cladding tube manufactured from French sponge 
Zr by the Russian process 

E11 0g(3&) 

Zr-l%Nb 

as-received El 10 cladding tube manufactured from 70% French 
sponge Zr, 30% iodide Zr and recycled scrap by the Russian proc¬ 
ess 

El 10o(3ru) 

Zr-1 %Nb 

as-received El 10 cladding tube manufactured from 70% Russian 
sponge Zr, 30% iodide Zr and recycled scrap 

E635 G(fr) 

Zr-1 %Nb-1.2%Sn- 
0.35%Fe 

as-received E635 cladding tube manufactured from French sponge 
Zr by the Russian process 


Table A-2. The geometry' of unirradiated cladding materials 


Parameter 

Cladding material 

(mm) 

E110 

E110A 

E110K 

El 10| OW Hf 

E635 

Zry-4 

E110o<fr) 

E 1 1 0o(3fr) 

El 10 G ,3ru) 

E635 G( fr) 

Outer 

diameter 

9.145 

9.148 

9.130 

9.170 

9.100 

9.142 

10.750 

9.140 

9.141 

9.130 

9.140 

Inner 

diameter 

7.73 

7.73 

7.72 

7.73 

7.78 

9.35 

7.74 

7.74 

7.73 

7.74 

Cladding 

thickness 

0.708 

0.709 

0.700 

0.725 

0.685 

0.681 

0.700 

0.700 

0.700 

0.700 

0.700 


Table A-3. Initial characteristics of irradiated cladding material 


Characteristic 

Unit 

Value 

The type of fuel assembly 

— 

VVER-1000 (the first unit of Zaporozhie 
Nuclear Power Plant) 

Fuel assembly number 

— 

E0325 

Fuel elements number 

— 

#156, #273 

Axial coordinates of fuel element section used 
for the refabrication of irradiated cladding 

mm 

• Fuel element #156: 1640-2890 

• Fuel element #273: 1800-2240 

Fuel bumup as a function of axial coordinate 

— 

see Fig. A-4 

Outer diameter of fuel element cladding as a 
function of axial coordinate 

— 

see Fig. A-5 

Average fuel bumup in the section of fuel 
element 

MWd/kg U 

• Fuel element #156: 52.0 

• Fuel element #273: 49.5 

Average outer cladding diameter in the section 
of fuel element 

mm 

• Fuel element #156: 9.03 

• Fuel element #273: 9.04 

Average cladding thickness in the section of 
fuel element 

mm 

• Fuel element #156: 0.68 

• Fuel element #273: 0.69 

Microstructure of irradiated cladding 

— 

see Fig. A-6 


A-5 
















































Characteristic 

Unit 

Value 

Average Zr0 2 thickness on the outer surface of 
irradiated cladding 

pm 

• Fuel elements ##156, 273: 5 

Average Zr0 2 thickness on the inner surface of 
irradiated cladding 

pm 

• Fuel elements ##156, 273: 0 

Hydrogen content in the irradiated cladding 

% by weight 

• Fuel elements ##156, 273: 0.0047 



Axial coordinate (mm) 


Fig. A-4. Axial burnup distribution for fuel element #156 and #273 (experimental data) 



Fig. A-5. Profilometry of cladding outer diameter as a function of axial coordinate 


A-6 
























































Fig. A-6. The microstructure of irradiated El 10 cladding before the oxidation tests 


A-4. Scoping tests performed to verify the experimental procedure of cladding temperature measurements 

The following scoping tests were developed and performed to validate and verify the procedure of temperature measure¬ 
ments at the oxidation of cladding samples: 

• scoping tests to reveal the axial temperature distribution inside the electric furnace; 

• scoping tests to reveal the temperature distribution as a function of cladding sample length; 

• scoping tests to obtain the comparative data characterizing two different methods of determination of cladding sample 
temperature: 

- the method on the basis of thermocouple, which was welded on the cladding surface; 

- the method on the basis of thermocouple, which was installed in the gas volume inside the cladding sample. 

The major results of these scoping tests are presented below. 

The data base characterizing the axial temperature distribution inside the electric furnace 

The procedure of this test consisted of the following options: 

• the electric furnace was heated to 1100 C (at the middle of furnace); 

• the thermocouple (which was installed along the axis of electric furnace) was moved the step by step from the posi¬ 
tion #1 to the #15 (Fig. A-7); 

• the processing of measured data allowed to obtain the axial temperature profile inside electric furnace; 

• these data were used to determine the “cold” and “hot” positions of cladding sample (see Fig. 3.2, section 3). 


A-7 

























Fig. A-7. The axial temperature distribution inside the electric furnace 


The measurement of axial temperature distribution along the cladding sample length 

These measurements were performed using the following procedure: 

• seven thermocouples were installed sequentially on the length of “hot” position (Fig. A-8); 

• the electric furnace was heated to 1100 C (at the “hot” position); 

• the processing of temperature measurements allow to reveal that the temperature nonuniformity is not more than 6 C. 



1200 


160 


<D 

g 1080 

D 


1040 


1000 


0 


20 





# 1-7 measured 
temperatures 

.1 .2. 

, 4 

^ 6 

7 














40 


60 


Axial coordinate (mm) 

Fig. A-8. The axial temperature profde at the “hot” position 


80 


A-8 




















































The develop ment ot comparative data to characterize two possible methods of measurement 

of cladding sample temperature 

The following experimental procedure was used to obtain the appropriate data: 

• the one thermocouple was welded on the outer surface of cladding sample; 

• the second thermocouple was installed inside cladding sample along the sample axis; 

• axial coordinates of both thermocouples were the same; 

• the cladding sample with thermocouples was installed in the “cold” position (see Fig. 3.2., section 3); 

• the electric furnace was heated to 1100 C; 

• the cladding sample was removed from the “cold" position to the “hot” position, after that the cladding sample was 
oxidized at the hold temperature (1100 C) during approximately 400 seconds and finally the cladding sample was re¬ 
moved from the “hot position to the cold position. Such a way, the oxidation of cladding sample at the F/F combina¬ 
tions of heating and cooling rates was performed. 

The results of these comparative temperature measurements are presented in Fig. A-9. 



Fig. A-9. The cladding and steam temperatures under oxidation conditions with F/F combinations 

of heating and cooling rates 


The analysis of obtained data allowed to make the following conclusions: 

• the differences in the both temperature measurements at the stationary mode do not exceed the instrumental error of 
thermocouples; 

• the thermocouple measuring the steam temperature under estimates the real cladding temperature at the heating mode, 
this thermocouple overestimates the cladding temperature at the beginning of cooling mode; 

• the appropriate calculations shown that the systematic error of determination of effective time of oxidation (caused by 
above mentioned transient effects under heating and cooling conditions) is very insignificant; 

• taking into account obtained comparative data it was decided to use the thermocouple placed inside the cladding sam¬ 
ple for the temperature measurement at the oxidation tests. 


A-9 




















A-5. The ECR measurement 


By definition, the equivalent cladding reacted (ECR) is a total thickness of zirconium layer that reacts with steam assum¬ 
ing that all absorbed oxygen were converted to a stoichiometric zirconia layer, divided by the initial thickness of the clad¬ 
ding. Besides it can be shown that the ECR is the ratio between the oxygen weight, which was absorbed by the cladding 
during the oxidation test to the oxygen weight, which could be absorbed by the cladding at the full (100 %) oxidation. 

To determine the experimental ECR, two alternative methods are used as a rule: 

• the measurement of oxygen weight gain using the determination of the cladding sample weight before and after the 
oxidation test; 

• the metallographic method which is based on the thickness measurement of the Zr0 2 and a-Zr(O) layers. 

However, the following specific features of the Zr-l%Nb (El 10) cladding material hamper the use of these methods: 

• the spallation and falling of oxide caused by the breakaway oxidation; 

• the nonuniform boundary between the a-Zr(O) layer and prior-P phase (that does not allow to measure the macro¬ 
scopic value of the a-Zr(O) layer using limited numbers of metallographic samples); 

• the inexact knowledge of the radial distribution of oxygen concentration in the a-Zr(O) and prior-P phase layers. 

Taking into account the above stated issues a special method of the ECR measurement was developed for this work. The 
general idea of this method is based on the following consideration: 

• the weight gain of the El 10 sample obtained during the oxidation test is determined as the difference between the 
weight gain of this sample oxidized up to 100% zirconium/niobium oxide (this value can be calculated with a high 
accuracy) and the weight gain of this sample at the additional full oxidation of the El 10 sample metallic part (re¬ 
mained after the test oxidation). 

In accordance with this approach, the following additional procedures were performed for the tested ring oxidized El 10 
sample: 

• the weight of the oxidized sample (as a rule, the weight of the ring oxidized sample fragments after the ring compres¬ 
sion tests) was measured; 

• the oxidized sample (or the sample fragments) was additionally oxidized at 1100 C in air atmosphere during 4 hours; 

• the weight of the completely oxidized (100 % oxide) sample was measured. 

This procedure (the extra oxidation of the El 10 ring samples) is explained in detail in Table A-4. 


Table A-4. The major provisions of the ECR determination procedure 


Stage of procedure 

Comments 

1. The weight gain of the El 10 cladding sample after the oxida¬ 
tion test (Am 0 ): 

Am 0 = AM ioo - Am c 

where AM )0 o - the weight gain of the cladding sample at the 
complete 100 % oxidation (g); 

Am c - the weight gain of the cladding sample, ob¬ 
tained during the extra oxidation procedure 
(g) 


2. The weight gain of the El 10 cladding sample at the complete 
oxidation in air atmosphere can be determined using the following 
equation: 

A M m = m[0.99 2A ° +0.01 5 A ° 1 = 0.3516 M 

v A Zr 2 A m j 

where A 0 = 16 (oxygen atomic mass); 

The following assumptions were used: 

• Zr concentration in the El 10 alloy is 0.99 (by 
weight); 

• the stoichiometric Zr0 2 and Nb 2 0 5 are formed 
at the complete oxidation of the El 10 alloy in 
the air 


A-10 










Stage of procedure 

Comments 

A Zr =91.22 (zirconium atomic mass); 

A Nb =92.91 (niobium atomic mass); 

M = the weight of the unoxidized ring sample (g) 


3. Thus, the ECR (%) can be calculated using the following equa¬ 
tion: 

ECR = ^ m " 100% = 1 

AM 100 0.3516 M 

Special scoping tests have been performed to estimate the calcula¬ 
tion accuracy on determining the AMioo parameter. Results of 
these tests have shown that the range of relative differences be¬ 
tween the calculated and measured AM I00 is 0.01-0.12 % 

It should be noted that the weight gain of the El 10 
cladding sample after the oxidation test (Amo) is 
not contained in the ECR equation. Therefore, the 
ECR measurement error is not the function of ox¬ 
ide weight and. consequently, the spallation and 
falling of the ZrO : oxide during the oxidation and 
handling (procedures after the test) are not impor¬ 
tant for the accuracy of the ECR measurement. 

This accuracy is determined by the accuracy of the 
sample weight measurements (the unoxidized sam¬ 
ple and the sample before and after the extra oxida¬ 
tion) and by the accuracy of several other meas¬ 
urements described below 

4. The unoxidized ring sample weight (M) can be determined us¬ 
ing the following considerations: 

M - m/-L 

where m/ - the weight of the unoxidized cladding sample 

1 cm long (g/cm); 

L - the unoxidized ring sample length (cm) 

Special scoping tests performed with nine cladding 
samples have shown that the w/ average value is 
0.01218 g/cm. The mean square error of this pa¬ 
rameter is very small (0.0001) 

5. The unoxidized ring sample length (L) is determined by 

L = L — 

°h 

where L 0 - the measured length of the oxidized ring sam¬ 
ple (cm); 

/ - the measured length of the unoxidized 
(100 mm) sample; 

l Q - the measured length of the oxidized (100 mm) 
sample 

The set of measurements allowed to reveal that the 
length (100 mm) of the El 10 sample is extended 
by 0.5-1 % during the oxidation 

6. The weight gain of the cladding sample obtained during the 
extra oxidation (Am f ) can be expressed by 

Am c — m c - m s 

where m s - the sample mass before the extra oxidation (g); 
m c - the sample mass after the extra oxidation (g) 


7. The sample mass before the extra oxidation (m s ) will be 

m s = m 0+ Kie + Mh 

where mo, Me - the mass of the sample metallic part and that 
of absorbed oxygen (g); 

m H - the mass of absorbed hydrogen (g) 

It should be taken into account that the increase in 
the El 10 sample mass is the result of two proc¬ 
esses: 

• oxygen absorption; 

• hydrogen absorption 

8. Thus, the ECR expression is 

f /S.m r m„ ^ , __ n/ 

ECR- 1 c -100% 

( 0.3516 M 0.3516 M) 


9. The absorbed hydrogen mass (m H ) can be determined by 
m H = Ch-ICT 6 m ( 



A-l 1 






















Stage of procedure 

Comments 

where C H - the measured hydrogen concentration in the 
ring sample (per-unit: ppm) 


10. Taking into account that the ratio of m s to M is close to 1, the 
hydrogen correction factor (AECR H = (m H /0.3516)100 %) can be 
presented with the reasonable accuracy by 

C -10 4 

A ECR,, = " 

0.3516 

C H (ppm) 10" 4 = C H (% by weight) 

11. Finally, the equation for the ECR determination has the fol¬ 
lowing form: 

ECR = 100%f 1 A '" c l ° ) C " 10 

v 0.3516 m, L 0 /) 0.3516 


12. The ECR measurement error was estimated using the follow¬ 
ing experimental data characterizing the errors of individual 
measurements: 


• the absolute error of m c measurement: 0.0002 g 

• the absolute error of / measurement: 0.01 cm 


• the absolute error of l Q measurement: 0.01 cm 

• the relative error of m t measurement: 1.6 % 

• the maximum relative error of Am c determination: 

- 2.3 % at the 5 % ECR 


- 1.2% at the 10% ECR 


• The relative error of L a measurement (using five azimutally 
distributed measurements): 0.8 % 


• The relative error of l/l 0 determination: 1 % 

The relative error of ECR measurement is 


• 3.1 % at the ECR=5 


• 2.4 % at the ECR=10 % 


13. The expression to determine the specific weight gain Am 0(S p) 
was formulated on the basis of the ECR data: 

The following equation was used to determine the 
specific weight gain: 

\ 0.3516 m, 1 10 at the double-sided 

A m ntsp\ ~ tLK . 

n(d Q + d ,) l Q 100% oxidation 

A m = Am ° 

LArrl 0(SP ) £ 

^ _ 0.3516 m, l 10“ ecr at the single-sided 

(HSP) 7Td 0 / 0 100% oxidation 

where Amo ( sp) - the specific weight gain (mg/cm - ); 

d 0 - the outer cladding diameter (cm); 
d, - the inner cladding diameter (cm) 

where: S - the sample surface area 

14. Several modifications were developed to determine the ECR 
in the irradiated claddings: 

• the weight of the irradiated cladding 1 cm long before the 
oxidation (mi) was determined 

• the initial ECR of irradiated claddings (the ECR developed 
during the irradiation) was determined (ECR, = 0.5 %) 

a) Several scoping tests were performed with the 
irradiated cladding samples to measure the parame¬ 
ter nii in accordance with the approach presented in 
the item 4. 

b) The following expression was used to determine 
the ECRj: 


A-12 




















Stage of procedure 

Comments 


Sfa 

ECR = 2 100% 

1.56 8 cl 

where &zroi - ZrO : layer thickness on the outer 
and inner cladding surfaces be¬ 
fore the oxidation (5 pm) 

5 c i - the irradiated cladding thickness 
(685 pm) 

1.56 - Pilling-Bedworth coefficient 

15. The final equation for the ECR in the irradiated cladding is 

ECR = 100%f 1 An ’ c ! ° 1 C " 10 +0.5% 

^ 0.3516 m, L 0 1) 0.3516 



A-6. Determination of the oxidation equivalent time 

It is known that to estimate the oxidation kinetics of the cladding, the experimental data characterizing the specific weight 
gain dependence on the oxidation time at the given temperature are required. But the real experimental oxidation history 
of the tested cladding sample consists of three test modes (Fig. A-10): 

1. The transient mode (the oxidation temperature transient) on the sample heating. 

2. The steady state mode (the oxidation temperature is approximately constant). 

3. The transient mode on the sample cooling. 



Fig. A-10. The oxidation history of the E110 as-received tube at 1100 C and F/F combination 

of heating and cooling rates 


A-13 


















































To transform the transient oxidation conditions into the oxidation kinetics (T=const), the equivalent time (te q ) is used: 


'ox 

f 


exp 


Q 




RT(t) 


dt 


r 


exp 


_Q_ 

R T 


eq y 


where: to X - time for the complete oxidation (s); 

Q - the activation energy (cal/mole); 

R - the gas constant (1.987 cal/mole K); 

T(t) - the oxidation temperature as a function of time (K); 

T eq - the assigned steady temperature (K). 

In accordance with the published data [1,2, 3], the following value for the activation energy was used: 
Q=39940 cal/mole. 


The appropriate estimates have shown that the relative error of the equivalent time on using other values for the activation 
energy do not exceed 0.3 %. 


References 


[1] Cathcart J.V., Pawel R.E. at. al. "Zirconium Metal-Water Oxidation Kinetics: IV. Reaction Rate Studies". 
ORNL/NUREG-17, 1977. 

[2] Bibilashvili Yu.K., Sokolov N.B., Salatov A.V., Andreyeva-Andriyevskaya L.N., Nechayeva O.A., Vlasov F.Yu. 
"RAPTA-5 Code: Modelling of Behaviour of Fuel Elements of VVER Type in Design Accidents. Verification Calcu¬ 
lations", Proc. of IAEA Technical Committee Meeting on "Behaviour of L WR Core Materials under Accident Condi¬ 
tions", Dimitrovgrad, Russia, on 9-13 October 1995. IAEA-TECDOC-921, Vienna, 1996. 

[3] Bohmert J., Dietrich M., Linek J. "Comparative Studies on High-Temperature Corrosion of Zr-l%Nb and Zircaloy-4", 

Nuclear Engineering and Design, 147 Nol, 1993. 


A-14 







APPENDIX B 

Tables with Results of Oxidation and Mechanical Tests: 

E110, E110A, E110K, EllOpol, E635, E110 G<fr) , E110 G(3ru} , E110 G(3fr) , 
El lO Um , H f E635 G( f r) , Zry-4 as-Received Tubes and El 10 Commercial 

Irradiated Claddings 








Table B-l. A summary list of tested unirradiated tube samples. Characterization of test conditions 


and major test results 


Material 

Oxidation 

type 

Heating/ 

Cooling 

comb. 

Steady 

temp. 

(C) 

Sample 

number 

Equiv. 

Time 

(s) 

Measured 

ECR 

(%) 

Weight gam 
(mg/cnT) 

Residual 

ductility 

(%) 

Relative 

displacement 

(%) 

H cont. 
(ppm) 

E110 

double¬ 

sided 

F/F 

800 

123 

7205 

3.4 

2.7 

59.8 

65.3 

80 

132 

28814 

10.7 

8.5 

8.0 

15.6 

466 

144 

21850 

8.6 

6.7 

16.0 

23.0 

150 

900 

130 

2407 

3.9 

3.1 

57.2 

58.2 

106 

131 

4804 

6.7 

5.4 

31.8 

32.1 

194 

142 

8120 

12.3 

9.6 

0.0 

2.7 

3010 

950 

140 

5012 

13.4 

10.5 

0.0 

3.2 

2780 

141 

2490 

11.2 

8.8 

0.0 

2.9 

920 

1000 

44 

865 

7.7 

6.2 

8.4 

13.0 


45 

798 

7.6 

6.1 

20.7 

24.4 


119 

871 

5.7 

4.5 

19.4 

26.4 

173 

1100 

28 

529 

10.5 

8.4 

0.6 

6.9 


30 

426 

8.9 

7.1 

0.2 

5.1 

1110 

41 

284 

8.2 

6.6 

0.0 

4.5 

1130 

46 

194 

6.5 

5.2 

57.2 

57.7 

75 

47 

209 

7.0 

5.6 

57.7 

58.3 

30 

62 

445 

8.0 

6.4 

4.9 

11.9 

660 

65 

488 

7.5 

6.0 

42.1 

44.3 

40 

68 

605 

10.0 

7.9 

3.3 

10.0 


81 

496 

8.1 

6.5 

1.5 

8.5 

280 

82 

499 

7.9 

6.3 

1.9 

6.9 

390 

83 

341 

10.0 

8.0 

0.5 

5.5 


85 

307 

6.0 

4.8 

* 

* 


86 

336 

6.9 

5.5 

_ * 

- * 


92 

555 

8.5 

6.8 

1.4 

7.2 

690 

96 

762 

9.8 

7.8 

_ * 

_ * 


104 

1814 

16.1 

12.8 

0.0 

4.8 


105 

818 

11.8 

9.4 

_ * 

- * 


110 

1714 

13.7 

10.7 

0.0 

4.4 


111 

770 

8.5 

6.7 

8.4 

14.2 


1120 

25 

403 

9.9 

7.9 

0.1 

4.7 

1010 

1200 

102 

364 

13.3 

10.6 

0.0 

4.5 

926 

106 

308 

11.5 

9.1 

0.0 

4.4 

430 

F/Q 

1100 

63 

250 

6.1 

4.9 

57.7 

58.2 

30 

66 

375 

7.1 

5.7 

47.3 

49.3 

90 

71 

445 

9.7 

7.7 

2.7 

9.4 

580 

F/S 

1100 

17 

** 

10.0 

8.0 

0.0 

4.2 

730 

19 

- 

14.2 

11.3 

0.0 

4.5 


35 

- 

11.0 

8.8 

0.0 

3.1 

1150 

37 

- 

9.4 

7.6 

14.1 

19.2 


40 

- 

8.7 

7.0 

6.1 

10.8 


48 

- 

8.7 

7.0 

56.2 

56.4 


52 

- 

6.2 

5.0 

57.0 

54.8 

180 

S/F 

1100 

27 

- 

4.6 

3.6 

40.3 

42.3 

150 

32 

- 

7.1 

5.7 

24.4 

28.0 


34 

- 

7.2 

5.8 

2.1 

8.5 


49 

- 

7.6 

6.1 

4.1 

11.2 


51 

- 

7.7 

6.2 

0.5 

5.6 



B-2 





























































Material 

Oxidation 

type 

Heating/ 

Cooling 

comb. 

Steady 

temp. 

(C) 

Sample 

number 

Equiv. 

Time 

(s) 

Measured 

ECR 

(%) 

Weight gain 

2 

(mg/cm') 

Residual 

ductility 

(%) 

Relative 

displacement 

(%) 

H cont. 
(ppm) 

E110 

double¬ 

sided 

S/S 

1100 

29 

- 

5.8 

4.6 

45.2 

39.7 

320 

31 

- 

8.0 

6.4 

21.3 

23.3 


33 

- 

7.5 

6.0 

3.8 

10.0 

530 

36 

- 

11.7 

9.4 

0.2 

3.8 

1460 

38 

- 

7.6 

6.1 

4.6 

9.5 

360 

single¬ 

sided 

F/F 

1100 

42 

1374 

8.9 

13.1 

24.4 

32.9 

540 

54 

936 

6.6 

9.8 

62.6 

70.4 

650 

55 

1457 

9.5 

14.4 

21.3 

28.3 

300 

56 

1245 

8.6 

12.7 

11.7 

19.2 

360 

76 

1821 

11.2 

16.5 

0.0 

3.3 


E110A 

double¬ 

sided 

F/F 

1100 

117 

675 

9.0 

7.1 

0.0 

5.1 

554 

E110K 

double¬ 

sided 

F/F 

1100 

58 

488 

14.0 

11.4 

0.0 

3.1 


67 

480 

9.7 

7.9 

0.0 

3.1 

1630 

88 

559 

7.0 

5.7 

0.0 

2.7 


E1 1 0lo» H/ 

double¬ 

sided 

F/F 

1100 

94 

672 

9.2 

7.1 

38.0 

41.6 

17 

121 

1289 

11.4 

8.8 

2.6 

9.0 

426 

EllOpol 

double¬ 

sided 

F/F 

1000 

119 

871 

4.3 

3.4 

59.6 

60.0 

44 

1100 

92 

555 

8.0 

6.4 

46.3 

52.0 

28 

E635 

double¬ 

sided 

F/F 

1000 

127 

670 

5.3 

4.2 

57.4 

57.6 

98 

138 

1493 

9.4 

7.5 

0.0 

5.5 

400 

1100 

60 

- 

8.8 

7.0 

0.4 

6.3 


61 

- 

9.6 

7.6 

0.0 

4.8 


126 

289 

7.8 

6.2 

30.4 

32.6 

138 

134 

300 

6.9 

5.5 

12.4 

18.7 

107 

135 

460 

9.3 

7.4 

7.2 

13.8 

35 

El lOoofr) 

double¬ 

sided 

F/F 

1100 

99 

755 

11.5 

9.1 

15.0 

23.6 

13 

El 10o<3ni) 

double¬ 

sided 

F/F 

900 

137 

14400 

7.5 

5.9 

35.5 

38.3 

66 

1000 

98 

2519 

6.9 

5.4 

42.4 

46.0 

16 

101 

5028 

8.9 

7.0 

28.3 

31.1 

11 

1100 

95 

739 

11.6 

9.1 

15.0 

22.5 

4 

97 

1548 

16.7 

13.1 

11.4 

17.3 

17 

109 

1743 

18.0 

14.2 

3.6 

9.2 

101 

1200 

112 

932 

23.5 

18.5 

0.0 

4.9 

2200 

120 

168 

7.8 

6.1 

3.2 

8.0 

824 

E110c fr , 

double¬ 

sided 

F/F 

1100 

89 

558 

10.5 

8.3 

11.7 

18.7 

15 

90 

933 

13.0 

10.3 

17.1 

22.8 

48 

1000 

91 

2016 

6.5 

5.1 

51.9 

54.3 

17 

93 

5013 

8.5 

6.7 

0.9 

6.6 

12 

E635G, fr , 

double¬ 

sided 

F/F 

1100 

100 

749 

12.5 

9.8 

4.3 

11.0 

18 

108 

572 

11.0 

8.6 

0.0 

2.5 


Zry-4 

double- 

sided 

F/F 

1100 

64 

495 

11.5 

9.5 

14.3 

22.4 

34 

S/S 

1100 

43 

- 

11.3 

9.3 

4.8 

12.7 

37 

El 10 m 

double¬ 

sided 

F/F 

1100 

136 

604 

7.7 

6.1 

1.1 

8.9 

90 


* - Tube sample was used for the three-point bending test 

** . Equivalent time calculation procedure was not developed for slow heating and slow cooling test modes 
F/F, F Q, F S. S/F, S S characterize the heating/cooling combinations of the oxidation mode 
(F - fast. Q - quench. S - slow) 

ECR. weight gain, residual ductility, relative displacement are average values obtained due to the procedure 
of processing of individual measurements these parameters on several rings cutted from the tube sample 
(see Table B-2) 


B-3 
























































Table B-2. A summary data base on results of oxidation and mechanical tests with unirradiated 

cladding samples' 


Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp . 2 

(%) 

Resid. 

Duct . 3 

(%) 

4 

Other investigations 

17 

E110 

1100 

10.0 

z 15 

10.3 


20 

4.1 

0.0 



double-sided 

F/S 


4 



20 

4.3 

0.0 






5 






MG 





625 * 

9.7 


20 

4.3 

0.0 






8 


730 




H2 

19 

E110 

1100 

14.2 

2,5 

13.9 


20 

3.9 

0.0 



double-sided 

F/S 


4 



20 

5.6 

0.0 






620 * 



20 

3.7 

0.0 






7 






MG, HV 





825 * 

14.4 


20 

4.8 

0.0 


25 

E110 

1120 

9.9 

2 






MG, HV 


double-sided 

F/F 


3 






MG 





4 

9.4 


20 

4.5 

0.0 






5 



20 

4.4 

0.0 






6 

10.0 


20 

6.0 

0.2 






7 

10.2 


20 

3.9 

0.0 






8 



20 

4.7 

0.2 






9 


1010 




H2 





10 



135,200,300 



MTt 

27 

E110 

1100 

4.6 

1 


199 




H2 


double-sided 

S/F 


2 



20 

48.1 

45.6 






3 



135 

63.0 

54.1 






4 

4.5 


20 

47.4 

46.9 






5 






MG 





6 

4.6 


20 

21.7 

16.5 






8 


90 




H2 





9 






MG 





10 



20 

52.1 

52.0 






11 


162 




H2 

28 

E110 

1100 

10.5 

5 

10.5 


20 

6.5 

0.4 



double-sided 

F/F 


6 



135 

15.0 

6.2 






7 



20 

7.2 

0.8 


29 

E110 

1100 

5.8 

2 

5.4 


20 

38.3 

38.1 



double-sided 

S/S 


3 






MG 





4 


319 




H2 





5 

5.5 


20 

37.6 

37.5 






6 



135 

63.5 

60.4 






7 



135 

64.5 

60.5 






8 

6.4 


20 

35.3 

34.9 






9 



20 

47.4 

47.2 


30 

E110 

1100 

8.9 

2 

9.2 


20 

4.0 

0.0 



double-sided 

F/F 


3 


1113 




H2 





4 



135 

9.4 

1.8 






5 






MG 





6 



135 

9.0 

1.7 






7 

9.2 










8 

8.3 


20 

6.7 

0.5 






9 

8.4 










10 

9.3 


20 

4.5 

0.0 



B-4 































































Tube 

sample 

number 

Material. 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp. 2 

(%) 

Resid. 

Duct. 3 

(%) 

Other investigations’ 1 

31 

E110 

1100 

8.0 

2 

9.3 


20 

5.8 

0.4 



double-sided 

S/S 


6 






MG 





8 

7.6 


20 

34.2 

33.9 






9 

7.2 


20 

30.0 

29.7 


32 

E110 

1100 

7.1 

2 

6.6 


20 

7.6 

0.6 



double-sided 

S/F 


5 



20 

18.7 

11.7 






6 






MG 





8 

7.6 


20 

43.7 

43.6 






9 

7.2 


20 

41.8 

41.6 


33 

E110 

1100 

7.5 

2 

7.2 


20 

14.5 

8.6 



double-sided 

s/s 


4 



135 

63.2 

60.7 






5 


870 




H2 





6 






MG 





7 


213 




H2 





8 

7.4 


20 

6.8 

1.0 






9 

7.8 


20 

8.5 

1.7 


34 

E110 

1100 

7.2 

2 

7.3 


20 

8.7 

1.7 



double-sided 

S/F 


4 

7.2 



12.7 

4.1 






5 






MG 





8 

7.1 


20 

9.8 

2.5 






9 

7.3 


20 

6.9 

0.1 


35 

E110 

1100 

11.0 

2 

10.9 


20 

3.7 

0.0 



double-sided 

F/S 


3 



135 

4.7 

0.0 






4 


1150 




H2 





5 






MG 





8 

11.0 


20 

3.1 

0.0 






10 



20 

2.6 

0.0 


36 

E110 

1100 

11.7 

7 

12.8 


20 

3.2 

0.2 



double-sided 

S/S 


4 


1457 




H2 





5 






MG 





6 

11.8 


20 

3.9 

0.2 






7 



135 

6.6 

0.7 






8 



200 

57.4 

57.0 






9 



300 

62.3 

62.0 






10 

10.6 


20 

4.3 

0.3 


37 

E110 

1100 

9.4 

1 






MG 


double-sided 

F/S 


2 

8.9 


20 

44.7 

44.4 






4 

9.1 


20 

20.2 

14.3 






5 



135 

64.3 

60.3 






7 

9.7 


20 

17.1 

10.9 






9 

9.7 


20 

7.2 

0.6 






10 

9.8 


20 

6.6 

0.4 






11 






MG 

38 

E110 

1100 

7.6 

2 



20 

7.9 

1.5 



double-sided 

S/S 


3 



135 

60.9 

60.3 






4 


315 




H2 





5 






MG 





7 

7.6 


20 

10.8 

3.6 






8 



135 

59.5 

58.4 






9 


400 

20 



H2. Fr 





10 



20 

9.8 

3.1 



B-5 

































































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heal Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

•y 

disp.' 

(%) 

Resid. 

Duct/ 

(%) 

Other investigations 4 

40 

E110 

1100 

8.7 

“> 

7.9 


20 

17.5 

11.8 



double-sided 

F/S 


5 






MG 





6 

9.4 


20 

8.5 

1.9 






8 



135 

17.7 

10.5 






10 

8.9 


20 

6.4 

0.3 


41 

E1I0 

1100 

8.2 

3 



135 

67.5 

60.1 



double-sided 

F/F 


4 



20 

6.5 

0.0 

Fr.HV.SEM 





5 






MG 





6 

8.1 


20 

4.0 

0.0 






8 


1130 




H2 





10 

8.3 


20 

4.9 

0.0 


42 

E110 

1100 

8.9 

1 


519 




H2 


single-sided 

F/F 





20 

14.6 

3.9 






3 

8.3 


20 

28.8 

22 1 






4 

9.1 


20 

61.8 

55.3 






5 






MG 





6 

9.2 


20 

26.2 

21.3 






7 


555 




H2 

43 

Zry-4 

1100 

11.3 

1 


37 




H2 


double-sided 

S/S 


2 

9.9 


20 

16.5 

8.6 






3 






MG 





4 






MG 





5 






MG 





6 

11.4 


20 

12.4 

4.6 






7 






MG 





8 



135 

18.7 

11.4 






10 

12.7 


20 

7.9 

1.1 






11 


271 




H2 

44 

E110 

1000 

7.7 

2 

9.9 


20 

3.8 

0.0 



double-sided 

F/F 


5 






MG 





6 

6.9 


20 

19.6 

14.1 






10 

6.2 


20 

15.7 

11.2 


45 

E110 

1000 

7.6 

2 

7.5 


20 

3.8 

0.0 



double-sided 

F/F 


5 






MG 





6 

7.8 


20 

58.9 

58.3 






10 

7.4 


20 

10.4 

3.8 


46 

E110 

1100 

6.5 

2 

6.0 


20 

58.9 

57.9 



double-sided 

F/F 


3 


129 




H2 





4 


20 




H2 





5 






MG 





6 

6.8 


20 

57.2 

56.8 






7 

6.5 


135 

71.6 

61.2 






10 

6.7 


20 

57.1 

56.9 


47 

E110 

1100 

7.0 

2 

7.1 


20 

58.9 

58.6 



double-sided 

F/F 


4 



135 

72.4 

64.2 






5 






MG 





6 

7.3 


20 

60.2 

59.8 






7 


30 




H2 





8 



20.135 



MTt 





9 

6.3 










10 

7.2 


20 

55.9 

54.6 



B-6 


































































Tube 

sample 

number 

Material. 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp." 

(%) 

Resid. 

Duct.' 

(%) 

Other investigations 4 

48 

E110 

1100 

8.7 

2 

8.6 


20 

51.4 

51.2 



double-sided 

F/S 


5 






MG 





6 

8.8 


20 

58.4 

58.1 






7 

9.1 


135 

69.7 

62.6 






10 

8.3 


20 

59.8 

59.4 


49 

E110 

1100 

7.6 

2 

7.5 


20 

10.3 

3.6 



double-sided 

S/F 


5 






MG 





6 

7.8 


20 

14.3 

7.4 






7 



20.135 



MTt 





8 



200. 300 



MTt 





10 

7.4 


20 

8.9 

1.3 


51 

E110 

1100 

7.7 




20 

8.8 

1.5 



double-sided 

S/F 


5 






MG 





6 

7.8 


20 

5.2 

0.0 






7 

7.6 


135 

10.9 

2.5 






10 



20 

2.7 

0.0 


52 

E110 

1100 

6.2 

2 



20 

46.3 

45.4 



double-sided 

F/S 


4 


180 




H2 





5 






MG 





6 

6.6 


20 

58.6 

58.2 






7 

6.5 


135 

72.5 

63.9 






10 

5.5 


20 

59.4 

59.1 


54 

E110 

1100 

6.6 

2 

6.6 


20 

72.1 

63.7 



single-sided 

F/F 


3 

6.0 


135 

73.7 

63.2 






4 


650 




H2 





5 






MG 





6 

7.0 


20 

67.5 

60.5 






7 

6.8 


135 

64.9 

53.9 






10 

6.8 


20 

71.8 

63.6 


55 

E110 

1100 

9.5 

2 

8.7 


20 

57.6 

51.8 



single-sided 

F/F 


3 

9.6 


135 

67.5 

55.0 






4 


300 




H2 





6 

10.0 


20 

15.0 

7.3 






7 

10.0 


135 

72.5 

61.9 






9 

9.4 


20 

12.3 

4.8 


56 

E110 

1100 

8.6 

2 

8.8 


20 

28.5 

21.8 



single-sided 

F/F 


3 

8.3 


135 

63.1 

52.4 






4 


360 




H2 





6 

9.0 


20 

15.1 

6.9 






7 

8.5 


135 

73.8 

61.3 






9 

8.3 


20 

13.9 

6.4 


58 

E110K 

1100 

14.0 

2 

14.8 


20 

3.4 

0.0 



double-sided 

F/F 


6 



20 

3.2 

0.0 






9 

13.2 


20 

2.8 

0.0 


60 

E635 

1100 

8.8 

2 

8.9 


20 

6.2 

0.3 



double-sided 

F/F 


5 






MG 





6 

8.6 


20 

7.4 

0.9 





* 

9 

8.9 


20 

5.2 

0.0 


61 

E635 

1100 

9.6 

0 

9.6 


20 

4.8 

0.0 



double-sided 

F/F 










B-7 
































































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp. 2 

(%) 

Resid. 

Duct.' 

(%) 

. . 4 

Other investigations 

62 

El 10 

1100 

8.0 

2 

7.6 


20 

13.7 

6.0 



double-sided 

F/F 


3 

8.1 


135 

62.8 

61.8 






4 


660 




H2 





6 

7.8 


20 

13.9 

7.1 






8 

8.3 


135 

31.9 

24.9 






9 

8.1 


20 

8.0 

1.5 


63 

E110 

1100 

6.1 

2 

5.5 


20 

60.6 

59.7 



double-sided 

F/Q 


3 

5.8 


135 

75.5 

65.4 






4 


30 




H2 





6 

6.1 


20 

58.9 

58.3 






8 

6.7 


135 

74.3 

62.8 






9 

6.6 


20 

55.2 

55.0 


64 

Zry-4 

1100 

11.5 

2 

11.3 


20 

21.7 

12.5 



double-sided 

F/F 


3 

11.4 


135 

26.5 

17.0 






4 


34 




H2 





5 






MG 





6 

11.7 


20 

21.9 

12.4 






8 

11.5 


135 

28.2 

18.8 






9 

11.8 


20 

23.6 

12.8 


65 

E1I0 

1100 

7.5 

2 

7.1 


20 

55.1 

54.9 



double-sided 

F/F 


3 

7.7 


135 

62.5 

59.2 






4 


40 




H2 





5 






MG 





6 

7.8 


20 

49.8 

49.5 






8 

7.4 


135 

72.0 

64.5 






9 

7.7 


20 

28.0 

21.9 


66 

El 10 

1100 

7.1 

2 

6.9 


20 

41.2 

35.9 



double-sided 

F/Q 


3 

7.0 


135 

67.7 

63.0 






4 


90 




H2 





5 






MG 





6 

7.1 


20 

57.2 

56.8 






8 

7.4 


135 

71.1 

62.4 






9 

7.1 


20 

49.4 

49.2 


67 

E110K. 

1100 

9.7 

2 

9.8 


20 

3.3 

0.0 



double-sided 

F/F 


3 

9.3 


135 

3.8 

0.0 






4 


1630 




H2 





5 






MG 





6 

8.7 


20 

3.1 

0.0 






8 

10.5 


135 

3.9 

0.0 






9 

10.4 


20 

2.9 

0.0 


68 

E110 

1100 

10.0 

2 

9.5 


20 

12.4 

4.7 



double-sided 

F/F 


3 

9.6 


135 

30.4 

22.6 






4 



200 

76.3 

65.0 






5 






MG 





6 

10.2 


20 

10.7 

4.1 






7 



300 

76.5 

65.5 






8 

10.0 


135 

29.1 

21.7 






9 

10.5 


20 

6.9 

1.0 



B-8 





























































Tube 

sample 

number 

Material. 
Oxid. Type 

Steady temp. 
(C) 

Heat Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H eonL 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

dispT 

(%) 

Resid. 

Duct.' 

(%) 

Other investigations’* 

71 

E110 

1100 

9.7 

*) 

9.1 


20 

17.4 

8.1 



double-sided 

FQ 


3 

9.9 


135 

55.7 

53.0 






6 

9.0 


20 

4.1 

0.0 






7 


581 




H2 





8 

10.1 


135 

52.1 

49.6 






9 

10.2 


20 

6.7 

0.0 


76 

E110 

1100 

11.2 

2 

11.9 


20 

3.2 

0.0 



single-sided 

FT 


3 



135 

5.6 

0.7 






6 

11.1 


20 

3.4 

0.0 






8 

10.0 


135 

12.2 

2.1 






9 



20 

3.3 

0.0 


81 

E110 

1100 

8.1 

1 








double-sided 

FT 



7.9 


20 

9.4 

2.3 






3 

8.3 


135 

25.1 

19.1 






4 


140 




H2 





5 






MG 





6 



20 

7.5 

0.8 






9 



20 

8.5 

1.4 






10 


422 




H2 

82 

E110 

1100 

7.9 




20 

7.1 

1 *> 



double-sided 

F/T 


3 

7.7 


135 

65.0 

64.2 






4 


170 




H2 





5 






MG 





6 



20 

7.3 







8 

8.0 


135 

66.1 

65.8 






9 



20 

6.2 

1.4 






10 


607 




H2 

83 

E110 

1100 

10.0 

1 

10.0 


20 

7.0 

0.9 



double-sided 

FT 


3 



20 

4.0 

0.0 


85 

E110 

1100 

6.0 




20 



3B 


double-sided 

FT 









86 

E110 

1100 

6.9 




20 



3B 


double-sided 

FT 









88 

E110K 

1100 


3 

7.0 


20 

2.7 

0.0 



double-sided 

FT 


4 



20 

2.6 

0.0 


89 

El lO&fr) 

1100 

10.5 

3 



135 

59.8 

56.5 



double-sided 

FT 


4 






MG 





5 



20 

21.2 

12.4 






6 

10.5 


20 

16.2 

10.9 






7 


7 




H2 





8 






H2 

90 

El 10(3(6) 

1100 

13.0 

3 



135 

60.1 

59.6 



double-sided 

FT 


4 






MG 





5 



20 

24.3 

18.2 






6 

13.0 


20 

21.4 

16.0 






7 


30 




H2 





8 


107 




H2 





10 


6 




H2 


B-9 




























































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mcch. Tests 
(C) 

Relative 

disp." 

(%) 

Resid. 

Duct.’ 

(%) 

Other investigations' 

91 

E * 10 G( fr) 

1000 

6.5 

3 



135 

60.8 

59.7 



double-sided 

F/F 


4 






MG 





5 

6.5 


20 

49.2 

48.8 






6 



20 

59.4 

55.0 






7 


5 




H2 





8 


28 




H2 

92 

EllOpol 

1100 

8.0 

2 






MG 


double-sided 

F/F 


3 


28 




H2 





4 



20 

52.0 

46.3 



E110 

1100 

8.5 

7 



20 

7.2 

1.4 



double-sided 

F/F 


8 


694 




H2 





9 

8.5 





MG 

93 

E1 1 Ocr(fr) 

1000 

8.5 

3 



135 

58.4 

53.8 



double-sided 

F/F 


4 






MG 





5 

8.5 


20 

6.3 

0.9 






6 



20 

6.8 

1.0 






7 


12 




H2 

94 

E110 lowHf 

1100 

9.2 

3 



135 

60.2 

58.5 



double-sided 

F/F 


4 






MG 





5 

9.2 


20 

45.2 

39.3 






6 



20 

37.9 

37.7 






7 


17 




H2 

95 

El 10(3<3ru) 

1100 

11.6 

1 

11.2 







double-sided 

F/F 


3 



135 

59.1 

57.4 






4 






MG 





5 



20 

22.6 

15.5 






6 

11.9 


20 

22.3 

14.4 






7 






MG 





8 


4 




H2 

96 

E110 

1100 

9.8 




20 



3B 


double-sided 

F/F 









97 

El 10 G(3ru) 

1100 

16.7 

1 

15.8 


20 





double-sided 

F/F 


3 



135 

53.8 

50.4 






4 






MG 





5 



20 

16.9 

9.6 






6 

17.0 


20 

25.9 

20.7 






7 

17.3 


20 

9.2 

4.0 






8 


17 




SEM, H2 





10 






MG 

98 

E1 1 0 G (3m) 

1000 

6.9 

3 



135 

60.1 

59.0 



double-sided 

F/F 


4 






MG 





5 



20 

44.0 

43.7 






6 

6.9 


20 

47.0 

41.0 






7 


16 




H2 

99 

El 10 C (3 fr) 

1100 

11.5 

3 



135 


57.1 



double-sided 

F/F 


4 






MG 





5 



20 

23.8 

16.0 






6 

11.5 


20 

23.4 

13.9 






7 


13 




H2 


B-10 






























































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H corn, 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp. 2 

(%) 

Resid. 

Duct. 3 

(%) 

Other investigations 4 

100 

E635G( fr ) 

1100 

12.5 

3 



135 

60.7 

59.5 



double-sided 

F/F 


4 






MG 





5 



20 

12.1 

4.1 






6 

12.5 


20 

9.9 

4.5 






7 


18 




H2 

101 

El lOo^n,, 

1000 

8.9 

3 



135 

55.0 

50.4 



double-sided 

F/F 


4 






MG 





5 



20 

28.3 

28.3 






6 

8.9 


20 

33.8 

28.3 






7 


11 




H2 

102 

E110 

1200 

13.3 

2 

13.0 


20 

4.6 

0.0 



double-sided 

F/F 


3 

13.5 


135 








4 






MG 





6 

13.3 


20 

4.5 

0.0 






7 


926 




H2 





8 



135 








9 



20 

4.5 

0.0 


104 

E110 

1100 

16.1 

5 






MG 


double-sided 

F/F 


6 

16.0 


20 

4.8 

0.0 






7 

16.0 










8 

16.2 


135 

4.2 

0.0 


105 

E110 

1100 

11.8 




20 



3B 


double-sided 

F/F 









106 

E110 

1200 

11.5 

2 



20 

4.4 

0.0 



double-sided 

F/F 


3 

11.4 


135 

7.7 

0.9 






4 






MG 





6 

11.5 


20 

4.4 

0.0 






7 


430 




H2 





8 



135 

8.9 

1.2 






9 



20 

4.3 

0.0 


108 

£635(5,^ 

1100 

11.0 

3 



135 

25.1 

19.6 



double-sided 

F/F 


4 






MG 





6 



20 

2.5 

0.0 






8 



135 

5.4 

0.0 


109 

E 1 1 0c(3ni) 

1100 

18.0 

2 

16.6 


20 

10.6 

5.4 



double-sided 

F/F 


3 

17.2 


135 

44.7 

44.3 






5 






MG 





6 

18.4 


20 

8.0 

1.9 






8 

19.0 


135 

29.3 

29.0 






9 

19.0 


20 

9.0 

3.6 






10 


101 




H2 

110 

E110 

1100 

13.7 

2 

13.5 


20 

3.8 

0.0 



double-sided 

F/F 


3 

13.5 


135 

19.2 

12.4 






5 






MG 





6 

14.0 


20 

3.9 

0.0 






8 

13.8 


135 

8.9 

2.5 






9 

13.8 


20 

5.6 

0.0 






10 


509 

20 

5.6 

0.0 

H2 


B-l 1 




























































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp." 

(%) 

Resid. 

Duct.’ 

(%) 

4 

Other investigations 

111 

E110 

1100 

8.5 

2 

7.9 


20 

4.4 

0.0 



double-sided 

F/F 


3 



135 

59.3 

58.5 






5 






MG 





6 

8.3 


20 

19.1 

12.3 






8 

8.8 


135 

59.5 

57.8 






9 

9.0 


20 

19.2 

13.0 


112 

HI 10^™, 

1200 

23.5 

2 






MG 


double-sided 

F/F 


3 

22.3 


20 

4.9 

0.0 






5 

24.7 


135 

3.2 

0.0 






6 


2200 




H2 

117 

E110A 

1100 

9.0 

3 

9.3 


135 

11.8 

4.7 



double-sided 

F/F 


5 






MG 





6 

8.1 


20 

5.1 

0.0 






8 

9.5 


135 

11.3 

4.2 






10 


554 




H2 

119 

El lOpol 

1000 

4.3 

2 






MG 


double-sided 

F/F 


3 

4.3 


20 

60.0 

59.6 






5 


44 




H2 


E110 

1000 

5.7 

8 

5.7 


20 

26.4 

19.4 



double-sided 

F/F 


9 


173 




H2 





10 






MG 

120 

H 1 1 0(j(3n,) 

1200 

7.8 

2 

8.1 


20 

18.3 

9.5 



double-sided 

F/F 


3 



135 

2.2 

0.0 






4 


824 




H2 





5 






MG 





6 

7.5 


20 

2.3 

0.0 






8 



135 

27.3 

20.5 






9 



20 

3.5 

0.0 


121 

Ell 0 lowHf 

1100 

11.4 

2 

12.1 


20 

4.5 

0.0 



double-sided 

F/F 


4 






MG 





6 

10.8 


20 

12.0 

5.0 






9 

11.3 


20 

10.4 

2.7 






10 


426 




H2 

123 

E110 

800 

3.4 

2 

3.0 


20 

67.7 

61.2 



double-sided 

F/F 


3 

3.0 


135 

67.4 

61.3 






5 






MG 





6 

3.5 


20 

63.8 

58.0 






7 


80 




H2 





8 

3.8 


135 

63.9 

56.9 






9 

3.5 


20 

64.5 

60.1 


126 

E635 

1100 

7.8 

2 

7.6 


20 

42.1 

41.9 



double-sided 

F/F 


3 

7.6 


135 

62.4 

62.0 






5 






MG 





6 

8.1 


20 

38.3 

37.9 






7 


138 




H2 





8 

7.9 


135 

56.9 

56.3 






9 



20 

17.5 

11.5 



B-12 





























































Tube 

sample 

number 

Material. 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

FI corn, 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp. 2 

(%) 

Resid. 

Duct. 3 

(%) 

Other investigations 4 

127 

E635 

1000 

5.3 

2 

4.8 


20 

60.1 

59.9 



double-sided 

F/F 


3 

4.8 


135 

70.0 

65.5 






5 






MG 





6 

5.7 


20 

54.7 

54.5 






7 


98 




H2 





8 

5.9 


135 

63.5 

61.2 






9 



20 

58.0 

57.7 


130 

E110 

900 

3.8 

2 



20 

57.3 

56.0 



double-sided 

F/F 


3 

3.7 


135 

64.0 

56.8 






5 






MG 





6 

3.8 


20 

60.4 

59.3 






7 


106 




H2 





8 



135 

64.9 

60.1 






9 



20 

56.8 

56.2 


131 

E110 

900 

7.4 

1 



20 

30.6 

30.2 



double-sided 

F/F 


3 



135 

63.4 

61.2 






5 






MG 





6 

7.4 


20 

28.4 

28.1 






7 


194 




H2 





8 



135 

59.0 

56.8 






9 



20 

37.2 

37.0 


132 

E110 

800 

11.0 

2 

9.4 


20 

23.6 

16.4 



double-sided 

F/F 


3 

9.7 


135 

70.4 

64.4 






S 






MG 





6 

11.7 


20 

12.3 

5.5 






7 


466 




H2 





8 

12.2 


135 

67.7 

59.0 






9 

11.9 


20 

10.9 

2.2 


134 

E635 

1100 

6.9 

2 

6.8 


20 

14.5 

8.7 



double-sided 

F/F 


4 






MG 





6 

7.0 


20 

15.5 

6.6 






7 


107 




H2 





9 

6.7 


20 

26.2 

21.9 


135 

E635 

1100 

9.3 

2 

8.7 


20 

16.3 

10.9 



double-sided 

F/F 


3 



135 

53.6 

52.5 






5 






MG 





6 

9.4 


20 

13.1 

5.4 






7 


35 




H2 





8 



135 

53.5 

52.1 






9 

9.8 


20 

12.0 

5.2 


136 

El 10m 

1100 

7.7 

7 

6.8 


20 

10.9 

3.3 



double-sided 

F/F 


3 



135 

61.9 

58.4 






5 






MG 





6 

8.0 


20 

7.7 

0.0 






7 


90 




H2 





8 



135 

59.0 

56.8 






9 

8.3 


20 

8.1 

0.0 



B-13 




























































Tube 

sample 

number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Average 

ECR 

(%) 

Ring 

sample 

number 

As- 

measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 

(C) 

Relative 

disp.“ 

(%) 

Resid. 

Duct.’ 

(%) 

• 4 

Other investigations 

137 

EH0 Wni) 

900 

7.5 

2 



20 

63.0 

61.3 



double-sided 

F/F 


3 

6.6 


135 

61.2 

59.8 






5 






MG 





6 

7.4 


20 

42.6 

41.9 






7 


66 




H2 





8 

7.9 


135 

60.1 

59.0 






9 

8.0 


20 

9.3 

3.3 


138 

E635 

1000 

9.4 

2 

8.8 


20 

6.9 

0.0 



double-sided 

F/F 


3 

8.7 


135 

12.0 

3.2 






5 






MG 





6 

9.9 


20 

3.7 

0.0 






7 


400 




H2 





8 

10.5 


135 

20.8 

13.1 






9 

9.0 


20 

6.0 

0.0 


140 

E110 

950 

13.4 

2 

13.7 


20 

2.9 

0.0 



double-sided 

F/F 


3 



135 

5.9 

0.0 






5 






MG 





6 

12.1 


20 

3.5 

0.0 






7 


2780 




H2 





8 



135 

6.4 

0.0 






9 

14.4 


20 

3.2 

0.0 


141 

El 10 

950 

11.2 

2 

11.2 


20 

3.5 

0.0 



double-sided 

F/F 


3 



135 

7.7 

0.0 






5 






MG 





6 



20 

3.1 

0.0 






7 


920 




H2 





8 



135 

5.3 

0.0 






9 



20 

2.2 

0.0 


142 

E110 

900 

12.3 

2 

10.0 


20 

3.3 

0.0 



double-sided 

F/F 


3 



135 

4.2 

0.0 






5 






MG 





6 

13.3 


20 

2.5 

0.0 






7 


3100 




H2 





8 



135 

4.0 

0.0 






9 

13.7 


20 

2.2 

0.0 


144 

E110 

800 

8.6 

5 






MG 


double-sided 

F/F 


6 

8.4 


20 

21.0 

15.0 






7 


150 




H2 





8 

8.8 


20 

25.0 

17.0 



11 Ring tensile tests and three point bending tests are indicated in "Other investigations" column. 


all other mechanical tests are ring compression tests 
'* Relative displacement at failure 
” Residual ductility at failure 
41 Abbreviations: 

MG metallographic cross-section 
HV microhardness measurements 
MTt ring tensile mechanical tests 
3B three point bending tests 
Fr fractography investigation 
SEM scanning electron microscope examinations 


B-14 




















































Table B-3. A summary list of tested irradiated commercial claddings . 


Characterization of test conditions and major test results 


Material 

Oxidation 

t>P e 

Heating/' 

Cooling 

comb. 

Steady 

temp. 

(C) 

Sample 

number 

Equiv. 

Time 

(s) 

Measured 

ECR 1 

(%) 

. 2 

Weight gain' 

T 

(mg/cm') 

Residual 

ductility 

(%) 

Relative 

displacement 

(%) 

H cont.' 
(ppm) 




1000 

16 

544 

6.3 

4.7 

18.7 

26.7 

470 

E110 

double- 

F/F 


17 

664 

8.6 

6.5 

37.0 

39.6 

250 


sided 



10 

232 

7.7 

5.8 

0.0 

8.4 

1690 




1100 

14 

312 

8.3 

6.2 

0.0 

8.3 

1410 





15 

327 

8.1 

6.1 

0.0 

6.6 

- 





20 

137 

6.3 

4.7 

26.1 

30.4 

30 





21 

223 

7.0 

5.2 

0.0 

6.3 

- 




1200 

18 

218 

16.0 

12.5 

0.0 

3.4 

- 



S/F 

1100 

3 

_* 

5.8 

4.3 

11.2 

20.4 

170 



S/S 

1100 

1 


8.5 

6.4 

3.2 

11.3 

280 





2 

_* 

10.5 

8.0 

1.9 

8.9 

270 


* - Equivalent time calculation procedure was not developed for slow heating and slow cooling test modes 
F F. F Q. F S. S T. S S characterize the heating/cooling combinations of the oxidation mode 
(F - fast. Q - quench. S - slow) 

A sum of the ECR before the oxidation (0.5 %) and the ECR during the oxidation 
' A sum of the weight gain before the oxidation (0.4 mg/cm") and the weight gain during the oxidation 
A sum of the H content before the oxidation (47 ppm) and the H content during the oxidation 


Table B-4. A summary' data base on results of oxidation and mechanical tests (ring compression tests) 

with irradiated commercial cladding samples 


Tube sample 
number 

Material. 
Oxid. Type 

Steady temp. 
(C) 

HeatyCool. 

Comb. 

Ring sample 
number 

As-measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 

(C) 

Relative 
disp. ! 
(%) 

Resid. 

Duct. 2 

(%) 

Other 

investigations 3 

1 

E110 

1100 

1 



20 

11.3 

3.2 



double- 

s/s 

2 



135 

30.5 

25.1 



sided 


3 


280 




H2 




4 






MG 




5 

8.5 






> 

E110 

1100 

1 



20 

8.9 

1.9 



double- 

S/S 

2 



135 

17.5 

10.9 



sided 


3 


270 




H2 




4 






MG 




5 

10.5 






3 

E110 

1100 

i 



20 

20.4 

11.2 



double- 

S/F 

2 



135 

67.6 

58.6 



sided 


3 


170 




H2 




4 






MG 




5 

5.8 






10 

E110 

1100 

1 



20 

8.4 

0.0 



double- 

F/F 

2 



135 

14.0 

2.0 



sided 


3 


1690 




H2 




4 






MG. HV 




5 

7.7 







B-15 






















































l ube sample 
number 

Material, 
Oxid. Type 

Steady temp. 
(C) 

Heat/Cool. 

Comb. 

Ring sample 
number 

As-measu¬ 
red ECR 

(%) 

H cont. 
(ppm) 

Temp, of 
mech. Tests 
(C) 

Relative 

disp.' 

(%) 

Resid. 

Duct.' 

(%) 

Other 

investigations 

14 

El 10 

1100 

1 



20 

8.3 

0.0 



double- 

F/F 

2 



135 

75.8 

68.7 



sided 


3 


1410 




H2 




4 






MG, HV 




5 

8.3 






15 

E110 

1100 

1 



20 

6.6 

0.0 



double- 

F/F 

2 



135 

72.5 

63.0 



sided 


3 










4 






MG, HV 




5 

8.1 






16 

El 10 

1000 

1 



20 

26.7 

18.7 



double- 

F/F 

2 



135 

48.0 

36.0 



sided 


3 


470 




H2 




4 






MG, HV 




5 

6.3 






17 

El 10 

1000 

1 



20 

39.6 

37.0 



double- 

F/F 

3 


250 




H2 


sided 


4 






MG, HV 




5 

8.6 






18 

E110 

1200 

1 



20 

3.4 

0.0 



double- 

F/F 

2 



135 

11.3 

1.2 



sided 


3 










4 






MG, HV 




5 

16.0 






20 

El 10 

1100 

1 



20 

30.4 

26.1 



double- 

F/F 

3 


30 




H2 


sided 


4 






MG 




5 

6.3 






21 

El 10 

1 100 

1 



20 

6.3 

0.0 



double- 

F/F 

4 






MG 


sided 


5 

7.0 







11 Relative displacement at failure 
2) Residual ductility at failure 
31 Abbreviations: 

MG - metallographic cross-section 
H2 - hydrogen content measurements 
HV - microhardness measurements 


Notice. The numeration of the rings begins from the bottom of the tube based on its location in the furnace. As a rule, parts o 
approximately 7 mm long were cut from top and bottom of oxidized tube and rejected to avoid end effects. After that the 8 mm ring: 
were cut and numbered from bottom. The photos of tubes in the report and in the appendices (the photos represent entire uncut tubes 
are shown such that the numeration increases from the right to the left. 


B-16 












































APPENDIX C 

Temperature Histories, Appearances and Microstructures of El 10 
Standard As-received Tubes after a Double-sided Oxidation at 1100 C 
and S/S, S/F, F/S Combinations of Heating and Cooling Rates 


c-i 








Fig. C-l. Typical temperature histories for cladding samples oxidized at S/S combination of heating 

and cooling rate 


C-2 





































































Fig. C-2. Typical temperature histories for cladding samples oxidized at S/F combination of heating 

and cooling rate 


C-3 
























































Temperature (C) Temperature (C) 



Fig. C-3. Typical temperature histories for cladding samples oxidized at F/S combination of heating 

and cooling rate 


C-4 
















































































ECR=11.7 % 


#36 



#31 


ECR=8.0 % 


ECR=7.6 % 


#38 



#29 



ECR=5.8 % 



Fig. C-4. Appearance of El 10 standard as-received samples as a function of the ECR after 
a double-sided oxidation at 1100 C and S/S combination of heating and cooling rat es 


ECR=7.7 % 


ECR=7.2 % 


#51 



#34 



#27 


ECR=4.6 % 



Fig. C-5. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1100 C and S/F combination of heating and cooling rat es 


ECR=11 % 


#35 




ECR=6.2 % 



Fig. C-6. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1100 C and F/S combination of heating and cooling rates 


C-5 




































Etched 


#29-3 

ECR=5.8% 


Etched 



100 uni 


srMr#S 


' Mi* 

['• \9-r 
■ ■■ —•ivTa'i 


wmc. 


S§2| if-VM&H 1 

*kz2!m$sm 


#38-5 

ECR=7.6% 



MM 




£y*K 


lOOjirn 


#36-5 

ECR=11.7% 



Etched 


1(H) Hm 


Polished 








**>«€ /» 



Fig. C-7. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

1100°C and S/S combination of heating and cooling rates 


C-6 

















































#27-5 

ECR=4.6% 


#34-5 

ECR=7.2% 


#51-5 

ECR-7.7% 


Etched 


Etched 


Etched 


Etched 


Etched 


Polished 




!>Oum 


Fig. C-8. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

1100°C and S/F combination of heating and cooling rates 


C-7 










































#52-5 

ECR=6.2% 


Etched 



Etched 


#40-5 

ECR=8.7% 



100 p.m 


#35-5 

ECR=11.0% 



Etched 


I IK) |am 


'>' V o' ' 


Polished 





Polished 






Fig. C-9. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

1100°C and F/S combination of heating and cooling rates 


C-8 





































APPENDIX D 

Temperature Histories, Appearances and Microstructures of El 10 
Standard As-received Tubes after a Double-sided Oxidation at 800, 900, 
950, 1000, 1100, 1200 C and F/F (F/Q) Combinations of Heating and 

Cooling Rates 


D-l 








Temperature (C) Temperature (C) 




Fig. D-l. Typical temperature histories for cladding samples oxidized at F/F combination of heating 

and cooling rate 


D-2 






























































1400 

1200 

1000 

u 

¥ 800 
I 600 

o 

H 

400 


200 

0 


0 


100 


1400 

1200 

1000 


u 


p 800 

= 

J— 

| 600 

3 

H 

400 

200 

0 


0 


100 



- 

-—- 

352 s 

-> 

#66 (ECR=7.1 %) 






T=1100 C 


/12 C/s 





j 23.4 C/s 










~>aa S Cic 


136.7 C/s 




_ OH.J Ls/S 













pech66 


200 


300 

Time (s) 


400 


500 


600 





387 s 


# 

71 (ECR=9.7 %) 






0.6 C^^ 

i r/c 







T= 1 100 C 







121.S C/s 







j 




172.5 C/s 



45.8 C/s 




















pech71 


200 


300 

Time (s) 


400 


500 


600 


Fig. D-2. Typical temperature histories for cladding samples oxidized at F/Q combination of 

heating and cooling rate 


D-3 




















































ECR=11% 



ECR=8.6% 


ECR=3.4% 


123+144+lS2.c<ir 




View A 

HI 32-5 


Fig. D-3. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 800 C and F/F combination of heating and cooling rates 


D-4 

















Etched 



Etched 



Etched 



Etched 




Etched 




Etched 




Fis. D-4. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

o 

800°C and F/F combination of heating and cooling rates 


D-5 
































ECR=12.3% 


ECR=6.7% 


ECR=3.9% 





142+131+130.ciir 



View A 


Fig. D-5. Appearance of El 10 standard as-received tubes as a function of the ECR after a 
double-sided oxidation at 900 C and F/F combination of heating and cooling rates 


D-6 





Etched 


#130-5 

ECR-3.9% 


Etched 



#131-5 

ECR=6.7% 


Etched 




Etched 



Etched 




Etched 




Fig. D-6. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

900°C and F/F combination of heating and cooling rates 


D-7 


































140+141 ciir 



View A 


Fig. D-7. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 950 C and F/F combination of heating and cooling rates 


D-8 






Etched 


#140-5 

ECR=13.4% 


Etched 


#141-5 
ECR=11.2% 



litOum 


Etched 



Etched 




Fig. D-8. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

950°C and F/F combination of heating and cooling rates 


D-9 































ECR=7.8% 


ECR=7.6% 


ECR=5.7% 



44+45+U9.cdr 





View A 


Fig. D-9. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1000 C and F/F combination of heating and cooling rates 


D-10 




Etched 


#119-2 

(as-received) 

ECR=5.7% 



Etched 


#45-5 

ECR=7.6% 


Etched 



IOO^m 


#44-5 

ECR-7.7% 



100 jim 


Etched 



Etched 




Fio. D-10. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

1000°C and F/F combination of heating and cooling rates 


D-l 1 


































#104 



ECR=10.5% 


#28 






104 + 110+2S+S2+65+46 cdr 


View A 


#110-4 




#65-10 



Fig. D-ll. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1100 C and F/F combination of heating and cooling rates 


D-12 

















Etched 



#65-5 ECR=7.6 % 



#41-5 ECR=8.2 % 


#30-5 ECR=8.9 % 




#68-5 ECR-10.0% 


#110-5 ECR-14.0% 



Fig. D-12. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 
1100 C and F/F combination of heating and cooling rates (ECR=6.5-14 %) 


D-13 




























. f '* 


v *k. 




i-T*** 




-C** . . ✓ f 




■Pri'A 


■ mm 


qtirV&F^ 

si*. 


t. "- , / v *,vc 

• V P '^vK'y - 4 {M ! v ;vS,V 


-T t/ / / I t 7 . '*^ -* 1 

is| 


Sample 

#46-5 

ECR-6.5 % 


#65-5 

ECR=7.6 % 


#81-5 

ECR=8.1 % 


#30-5 

ECR=8.9 % 


#68-5 

ECR-10% 


#110-5 

ECR-14% 


#104-5 
ECR=16 % 


External surface (Etched) 


Internal surface 


Fig. D-13. Microstructure of Zr0 2 and a-Zr(O) layers in E110 standard as-received tubes 
after a double-sided oxidation at 1100 C and F/F combination of heating and cooling rates 


D-14 


























































ECR=11.4% 



--- , - r^r-i •- i * —* -*+> ♦ • rrm -*r- - .^»^p *-**%♦» 




102+106+113 c Jr 




Men A 

HI 02-4 


HI 06-4 


Fig. D-14. Appearance of El 10 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1200 C and F/F combination of heating and cooling rates 


D-15 






#106-4 


ECR=11.5% 


#102-4 


Etched 


Etched 


ECR=13.3 % 





Etched 


Etched 



Fig. D-15. Microstructure of El 10 standard as-received tubes after a double-sided oxidation at 

1200°C and F/F combination of heating and cooling rates 


D-16 
































APPENDIX E 

Appearance and Microstructure of El 10 Standard As-received Tubes 
after a Single-sided Oxidation at 1100 C and F/F Combination 

of Heating and Cooling Rates 


E-l 






#76 


ECR=11.2% 


ECR=9.5% 


ECR=8.9% 


ECR=8.6% 


ECR=6.6% 







#55 


#42 


#56 


#54 


0 7 6+ 0.55 + 04 2+ 056+054 cdr 




View A 


#56-1 


Fig. E.l. Appearance of El 10 standard as-received tubes as a function of the ECR after single-sided 
oxidation at 1100 C and F/F combination of heating and cooling rates 


E-2 











#42-5 


#54-5 


ECR=6.6 


% 


ECR=8.9 % 



Fig. E.2. Microstructure of El 10 standard as-received tubes after a single-sided oxidation at 1100 C 

and F/F combination of heating and cooling rates 


E-3 















































APPENDIX F 

Appearances and Microstructures of E635 Standard As-received Tubes 
after a Double-sided Oxidation at 1000, 1100 C and F/F Combination 

of Heating and Cooling Rates 


F-l 





ECR-9.4% 


ECR=5.3% 


13S+127.cdr 




View A 



ft 12 7-5 



Fig. F-l. Appearance of E635 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1000 C and F/F combination of heating and cooling rates 


F-2 











ECR=9.3% 


#135 






135-1:6-134 cdr 


V i cm A 


U60-1 (ECR=8.8°o) 



#134-4 



1 rrm 


Fig. F-2. Appearance of E635 standard as-received tubes as a function of the ECR after 
a double-sided oxidation at 1100 C and F/F combination of heating and cooling rates 


F-3 


























Etched 


#127-5 

ECR-5.3% 



lOOnin 


Mil 






■7 


-yy- 


Etched 


#138-5 

ECR-9.4% 



100)nn 


' • L.* • 

K'ty* 


",••• *■•'■■■ > • , • • .v': 

\v . 

Or . 

SIR 


- v '-^ 

.<* vv> > iff, 


• :„ -i. 

4&s* 

Mk 


!3 

;*:<v 




Etched 






i r J 5 . 



Etched 



>> « ' 






Fig. F-3. Microstructure of E635 standard as-received tubes after a double-sided oxidation at 1000 C 

and F/F combination of heating and cooling rates 


F-4 










































Etched 


#134-4 

ECR=6.9% 


Etched 



#126-5 

ECR=7.8% 


Etched 



#135-5 

ECR=9.3% 



l(K)|im 


Etched 



Etched 





Fig. F-4. Microstructure of E635 standard as-received tubes after a double-sided oxidation at 1100 C 

and F/F combination of heating and cooling rates 


F-5 















































































































































































































APPENDIX G 

Appearances and Microstructures of Ziy-4 As-received Claddings after 
a Double-sided Oxidation at 1100 C and S/S, F/F Combinations 

of Heating and Cooling Rates 


G-l 





ECR=11.5% 

F/F 


#64 



#43 


ECR=11.3% 

S/S 



Fig. G-l. Appearance of Zry-4 as-received claddings after a double-sided oxidation at 1100 C 

and S/S, F/F combinations of heating and cooling rates 





Etched 


#43-5 ECR=11.3 % (S/S) 


. A- * 




,v *% 

' Vv • 

V- x ^ : 

x- f , 

. . _ . K» s • - * 


V s ', • f-*:.: ■ 


#64-5 ECR=11.5 % (F/F) 


Fig. G-2. Microstructure of Zry-4 as-received claddings after a double-sided oxidation at 1100 C 

and S/S, F/F combinations of heating and cooling rates 


G-2 













































APPENDIX H 

Appearance and Microstructure of E110, E635 As-received Tubes 
Manufactured on the Basis of the Sponge Zr, El 10i owH fAs-received 
Tubes after a Double-sided Oxidation at 900,1000, 1100, 1200 C and 
F/F Combination of Heating and Cooling Rates 


H-l 






View A 

# 93-2 



— ■ I mm 


093G+091G.c<lr 


Fig. H-l. Appearance of E110 G (fr) as-received tubes as a function of the ECR after 
a double-sided oxidation at 1000 C and F/F combination of heating and cooling rates 



137.cdr 


View A 

# 137-5 



Fig. H-2. Appearance of E110 G <3ru) as-received tubes as a function of the ECR after 
a double-sided oxidation at 900 C and F/F combination of heating and cooling rates 


H-2 





















101+09S cdr 




Fig. H-3. Appearance of EllOcorm as-received tubes as a function of the ECR after 
a double-sided oxidation at 1000 C and F/F combination of heating and cooling rates 



Fig. H-4. Microstructure of E110 G (3ru) as-received tubes after a double-sided oxidation at 900°C 
and ECR=7.5 % (F/F combination of heating and cooling rates) 


H-3 






















#98-4 

HCR 6.9% 



1 1 10 


U(t'r) 


Etched 


Etched 


Etched 


Etched 





Fig. H-5. Microstructure of EllOo^™* and EllO^fr) as-received tubes after a double-sided oxidation at 1000°C 

and ECR=6.5-6.9 % (F/F combination of heating and cooling rates) 


H-4 
























Fig. H-6. Microstructure of E110 G <3ru) and E110 G <fr) as-received tubes after a double-sided oxidation at 1000°C 

and ECR=8.5-8.9 % (F/F combination of heating and cooling rates) 


H-5 

































View A 

UH9-4 



-1 mm 


090+0S9 cdr 

Fig. H-7. Appearance of Ell0 G (fr) as-received tubes as a function of the ECR after a doubie-sided oxidation 

at 1100 C and F/F combination of heating and cooling rates 


H-6 













ECR=18.0% 


#109 (ElIO al J 





109+9^+95+99 c.ir 


View A 

HI 09-4 



-1 rrm 


Fig. H-8. Appearance of E 110 G( 3 ru) and E110 G <3f r ) standard as-received tubes as a function of the ECR 
after a double-sided oxidation at 1100 C and F/F combination of heating and cooling rates 


H-7 




































#89-4 

ECR=10.5% 
(El 10 G(fr) ) 


Etched 


#90-4 

ECR=13.0% 

(El 10cj(fr)) 


Etched 


Etched 


Etched 


Etched 


#99-4 

ECR=11.5% 

(El 10 G (3f r) ) 


Etched 


'■-V V ' v 


Fig. H-9. Microstructure of E110 G (« r ), E110 G (3fr) as-received tubes after a double-sided oxidation at 1100°C 

and F/F combination of heating and cooling rates 


H-8 



















































#95-7 

ECR=11.6% 


#97-10 

ECR-16.7% 


#109-5 

ECR=18.0% 


Etched 


Etched 


Etched 


Etched 


Etched 


Etched 


Fig. H-10. Microstructure of E110 G <3ru) as-received tubes after a double-sided oxidation at 1100°C 

and F/F combination of heating and cooling rates 


H-9 

































ECR=22.3% 


m2 





1D+I1S+ I20.cdr 




View A 
# 112-1 


U120-5 


Fig. H-ll. Appearance of E110 G (3ru) as-received tubes as a function of the ECR after a double-sided oxidation 

at 1200 C and F/F combination of heating and cooling rates 


H-10 








# 120-5 


# 112-2 


ECR=7.8 % 


ECR=23.5 % 


Etched 


Etched 



Etched 



Etched 



Fig. H-12. Microstructure of E110 G <3ru) as-received tubes after a double-sided oxidation at 1200°C 

and F/F combination of heating and cooling rates 


H-l 1 





























ECR=12.5% 

#108 



100+ lOS.edr 



View A 

# 100-2 



■ ■ ■ I nun 


#108-4 


ECR-11.0% 


#100-4 


ECR=12.5 % 


Etched 


Etched 



100 |im 




•iW ■ .W 




■MigS 


Etched 



' ■ ■ . „ 

V. • • 

■ m & ' _ 

^• >; % .'riv? ■■ytd r ' -V;.. -5 

'■ . ■*? - ^ 

s SM 6 K 9 K teas 


d^MWt 


,. 

■*/' di K 


10( 1 j-tin 


Etched 



Fig. H-13. Appearance and microstructure of E635 C (f r ) as-received tubes after a double-sided oxidation at 1100°C 

and F/F combination of heating and cooling rates 


H-12 



















































Fig. H-14. Appearance of El 10 tow H f as-received tubes after a double-sided oxidation at 1100 C 

and F/F combination of heating and cooling rates 



Fia. H-15. Microstructure of El 10, ow H f as-received tubes after a double-sided oxidation at 1100 C 

o 

and F/F combination of heating and cooling rates 


H-13 































136.cdr 

Fig. H-16. Appearance of E110 m as-received machined/etched tubes after a double-sided oxidation at 1100 C 

and F/F combination of heating and cooling rates 


View A 


n 136-5 



■ - 



Fig. H-17. Microstructure of E110 m as-received machined/etched tubes after a double-sided oxidation at 1100 C 

and F/F combination of heating and cooling rates 


H-14 




















APPENDIX I 

Appearance and Microstructure of El 10 Commercial Irradiated 
Cladding after a Double-sided Oxidation at 1000,1100, 1200 C and S/S, 
S/F, F/F Combinations of Heating and Cooling Rates 


i-1 








ECR=7.7 % 


ECR=7.0 % 


#10 


#21 




ECR=6.3 % 


ECR=0.5 % 


#20 



before the test 



Fig. 1-1. Appearance of El 10 commercial irradiated claddings before and after a double-sided 
oxidation at 1100 C and F/F combination of heating and cooling rates 




ECR=5.8 % 
S/F 



Fig. 1-2. Appearance of El 10 commercial irradiated claddings after a double-sided oxidation at 
1100 C and S/S, S/F combinations of heating and cooling rates 


1-2 





















#18 


ECR=16.0 % 
1200 C 



#17 


ECR=6.5 % 
1000 C 



#16 


ECR=4.7 % 
1000 C 



Fig. 1-3. Appearance of El 10 commercial irradiated claddings after a double-sided oxidation at 
1000 C, 1200 C and F/F combination of heating and cooling rates 

















Etched 



Etched 


Etched 



Polished 


#1-4 

ECR=8.5% 

S/S 


Etched 



Polished 



100 nm 








Fig. 1-4. Microstructure of El 10 commercial irradiated cladding after a double-sided oxidation at 

1100 C and S/F, S/S combinations of heating and cooling rates 


1-4 









































































Fig. 1-5. Microstructure of El 10 commercial irradiated claddings before and after a double-sided 
oxidation at 1100 C and F/F combination of heating and cooling rates 


1-5 





































Sample 

Before 

oxidation tests 
(Polished) 


External surface 



Internal surface 


#20-4 

ECR=6.3 % 
(Etched) 




#21-4 

ECR=7.0 % 
(Etched) 


#10-4 

ECR=7.7 % 
(Polished) 



#15-4 

ECR=8.1 % 
(Polished) 


#14-4 

ECR=8.3 % 
(Polished) 





Fig. 1-6. Microstructure of Zr0 2 and a-Zr(O) layers in El 10 commercial irradiated cladding 
before and after a double-sided oxidation at 1100 C and F/F combination of heating and cooling 

rates 


1-6 














































































Etched 


#16-4 
ECR=4.7% 
T=1000 C 



#17-4 
ECR=6.5% 
T=1000 C 



Etched 


#18-4 

ECR=16.0% 
T=1200 C 



100|im 


Polished 



Polished 



Polished 



Fig. 1-7. Microstructure of El 10 commercial irradiated cladding after a double-sided oxidation at 
1000°C, 1200 C and F/F combination of heating and cooling rates 


1-7 









































































































NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 

(9-2004) 

NRCMD 3.7 

1. REPORT NUMBER 
(Assigned by NRC, Add Vol., Supp.. Rev., 
and Addendum Numbers, if any.) 

BIBLIOGRAPHIC DATA SHEET 


(See instructions on the reverse) 

NUREG/IA-0211 

2. TITLE AND SUBTITLE 

3. DATE REPORT PUBLISHED 

EXPERIMENTAL STUDY OF EMBRITTLEMENT OF 

ZR-1% NB WER CLADDING UNDER LOCA RELEVANT 

CONDITIONS 

MONTH YEAR 

March 2005 


4. FIN OR GRANT NUMBER 


Y6789 

5. AUTHOR(S) 

6. TYPE OF REPORT 

L. Yegorova, K. Lioutov, N. Jouravkova, A. Konobeev 

V. Smirnov, V. Chesanov, A. Goryachev 

IRSN 2005-194 NSI RRC Kl 3188 

7. PERIOD COVERED (Inclusive Dates) 


8 PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC. provide Division. Office or Region. U S. Nuclear Regulatory Commission, and mailing address: if contractor 
provide name and mailing address) 


Nuclear Safety Institute of Russian Research Centre "Kurchatov Institute" Moscow, Russian Federation 
State Research Centre "Research Institute of Atomic Reactors" Dimitrovgrad, Russian Federation 


9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC. type "Same as above': if contractor provide NRC Division. Office or Region. U S Nuclear Regulatory Commission, 
and mailing address.) 

Division of Systems Analysis and Regulatory Effectiveness 
Office of Nuclear Regulatory Research 
U.S. Nuclear Regulatory Commission 
Washington, DC 20555-0001 

10. SUPPLEMENTARY NOTES 


11. ABSTRACT (200 words or less) 

During 2001-2004, research was performed to develop test data on the embrittlement of niobium-bearing (Zr-1%Nb) cladding of 
the WER type under loss-of-coolant accident (LOCA) conditions. Procedures were developed and validated to determine the 
zero ductility threshold. Pre-test and post-test examinations included weight gain and hydrogen content measurements, 
preparation of metallographic samples, and examination of samples using optical microscopy, scanning electron microscopy 
and transmission electron microscopy. Sensitivity of the zero ductility threshold to heating and cooling rates was determined. 
Oxidation kinetics and ductility threshold were measured for the standard El 10 alloy, six variants with different impurities, 
Zircaloy, and irradiated El 10. Oxidation temperatures were varied from 800-1200 C, and mechanical (ring compression) testing 
temperatures were varied from 20-300 C. It was concluded that (a) the current type of El 10 cladding has an optimal 
microstructure, (b) oxidation and ductility of the oxidized cladding are very sensitive to microchemical composition and surface 
finish, (c) the use of sponge zirconium for fabrication of cladding tubes provides a significant reduction of the oxidation rate and 
an increase in the zero ductility threshold, and (d) additional improvement in oxidation and ductility can be achieved by polishing 
the cladding surface. 


12. KEY WORDS/DESCRIPTORS (Ust words 

or phrases that will assist researchers in locating the report.) 

13. AVAILABILITY STATEMENT 

Alloy 

Niobium 

unlimited 

Cladding 

Oxidation 

14 SECURITY CLASSIFICATION 

E110 

Ring Compression 

(This Page) 

Hydrogen Content 

Russian 

unclassified 

Kurchatov Institute 

Zirconium 

(This Report) 

Loss-of-Coolant Accident 
Material Properties 

Zircaloy 

unclassified 


15. NUMBER OF PAGES 

16. PRICE 


NRC FORM 335 (9-2004) 


PRINTED ON RECYCLED PAPER 






















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UNITED STATES 

NUCLEAR REGULATORY COMMISSION 

WASHINGTON, DC 20555-0001 


























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